Ralph J. Volino

United States Naval Academy, Annapolis, Maryland

Mounir B. Ibrahim

Cleveland State University, Cleveland, Ohio

Flow Control Under Low-Pressure Turbine 

Conditions Using Pulsed Jets

Final Report 

NASA/CR—2012-217415

September 2012

Notice for Copyrighted Information 

This manuscript has been authored under a NASA Interagency Agreement No. NNC07IA10I (through U.S. Navy Agreement 

No. N00189–07–P–A253). The United States Government has a nonexclusive, irrevocable, worldwide license to prepare 

derivative works, publish or reproduce this manuscript, and allow others to do so, for United States Government purposes. 

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 7115 Standard Drive

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Ralph J. Volino

United States Naval Academy, Annapolis, Maryland

Mounir B. Ibrahim

Cleveland State University, Cleveland, Ohio

Flow Control Under Low-Pressure Turbine 

Conditions Using Pulsed Jets

Final Report 

NASA/CR—2012-217415

September 2012

Notice for Copyrighted Information 

This manuscript has been authored under a NASA Interagency Agreement No. NNC07IA10I (through U.S. Navy Agreement 

No. N00189–07–P–A253). The United States Government has a nonexclusive, irrevocable, worldwide license to prepare 

derivative works, publish or reproduce this manuscript, and allow others to do so, for United States Government purposes. 

Any publisher accepting this manuscript for publication acknowledges that the United States Government retains such a 

license in any published form of this manuscript. All other rights are retained by the copyright owner. Attachments are 

reprinted by permission. 

National Aeronautics and

Space Administration

Glenn Research Center

Cleveland, Ohio 44135

Prepared under NASA Interagency Agreement No. NNC07IA10I 

(through U.S. Navy Agreement No. N00189–07–P–A253)
Acknowledgments

This research was made possible by a NASA Interagency Agreement No. NNC07IA10I (through U.S. Navy Agreement 

No. N00189–07–P–A253), awarded in response to NASA Research Announcement (NRA) solicitation (2006). The research was 

in support of the Subsonic Fixed-Wing Project of NASA’s Fundamental Aeronautics Program. The period of performance was 

January 2007 to September 2010. The following served as the Technical Offi cers (Technical Monitors or COTR’s): Anthony J. 

Strazisar, James D. Heidmann, and David E. Ashpis, NASA Glenn Research Center. The authors appreciated their support, as well 

as useful discussions and comments. American Society of Mechanical Engineers (ASME) permission to reproduce copyrighted 

material is greatly appreciated.

Available from

NASA Center for Aerospace Information

7115 Standard Drive

Hanover, MD 21076–1320

National Technical Information Service

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Available electronically at http://www.sti.nasa.gov

Trade names and trademarks are used in this report for identifi cation 

only. Their usage does not constitute an offi cial endorsement, 

either expressed or implied, by the National Aeronautics and 

Space Administration.

This work was sponsored by the Fundamental Aeronautics Program 

at the NASA Glenn Research Center.

Level of Review: This material has been technically reviewed by NASA technical management.
Flow Control Under Low-Pressure Turbine Conditions 

Using Pulsed Jets 

Ralph J. Volino 

United States Naval Academy 

Annapolis, Maryland 21402 

Mounir B. Ibrahim 

Cleveland State University 

Cleveland, Ohio 44115 

Summary 

In this study the effects of unsteady wakes and flow control using vortex generator jets (VGJs) were 

studied experimentally and computationally on the flow over a low pressure turbine airfoil. High loading 

is desirable for turbomachinery airfoils to reduce the number of airfoils and stages required to produce a 

given amount of power. Reducing the number of airfoils can reduce cost and weight and lead to increased 

efficiency. Increased loading, however, can lead to boundary layer separation as the magnitude of the 

adverse pressure gradients on the suction side of the airfoils is increased. Separation leads to loss of lift 

and increased aerodynamic losses, which reduce power and efficiency. The focus of the present study is 

on the low pressure turbine (LPT). The LPT is the heaviest component in an engine and produces the net 

output power. Any improvements to LPT weight and efficiency therefore have a large impact on overall 

engine performance. Boundary layer separation is a particular problem in the LPT because Reynolds 

numbers are typically low, particularly in aircraft engines at altitude, where the low density of the air can 

result in very low Reynolds numbers. The focus of the present work is avoiding or controlling separation 

while maintaining high loading. Work was done under simulated LPT conditions using a modern, very 

highly loaded airfoil. 

The airfoil selected for the study is the L1A, which was designed at the Air Force Research 

Laboratory. It is an aft loaded design with very high lift (Zweifel=1.35). Aft loading is desirable in that it 

reduces the strength of secondary flows, but it also requires a strong adverse pressure gradient on the 

downstream section of the airfoil, which makes the boundary layer more likely to separate. The L1A was 

designed specifically to provide a challenging case for separation control. In steady flow without flow 

control, the flow over the L1A is well behaved at high Reynolds numbers, but at low Reynolds numbers 

the boundary layer separates from the suction side of the airfoil and does not reattach. This results in a 

partial loss of lift and very high aerodynamic losses. 

An experimental facility consisting of a six passage linear cascade in a low speed wind tunnel was 

constructed and used in the experimental study at the U.S. Naval Academy. In the first phase of the work, 

the flow through the cascade was documented under steady inflow conditions without flow control. 

Reynolds numbers (based on suction surface length and nominal exit velocity from the cascade) ranged 

from 25,000 to 300,000. Inlet freestream turbulence levels of 0.6% and 4% were used. For Re≥200,000 

the separation bubble was small. For Re≤50,000 the boundary layer separated and did not reattach. The 

boundary layer was laminar at separation and transition to turbulence occurred in the separated shear 

layer in all cases. Transition did not cause reattachment at the low Reynolds numbers. This result 

contrasted with previous work on less aggressive airfoils, in which transition in the shear layer forced 

reattachment. With the L1A, the adverse pressure gradient is so strong that turbulence in the shear layer is 

not always able to force reattachment. At intermediate Reynolds numbers reattachment depended on the 

freestream turbulence. With high freestream turbulence the shear layer was more likely to reattach after a 

large separation bubble, while with low freestream turbulence it was more likely to remain separated. 

Large separation bubbles resulted in a large increase in profile loss by a factor of up to 5, a reduction in 

flow turning of as much as 16 degrees, and a reduction in lift of up to 20%. 

NASA/CR—2012-217415 1
In parallel with the experimental work, computational work was done at Cleveland State University. 

Reynolds Averaged Navier Stokes (RANS) calculations were done using a variety of turbulence models. 

With most models, transition was predicted too early, and separation was not predicted, even at low 

Reynolds numbers. Two models, the Standard k- model with shear stress transport (SKW-SST) and the 

four equation Transition shear stress transport model (Transition-SST) did better. Of these two, the 

Transition-SST model generally gave results in better agreement with the experiments. Over the full 

range of Reynolds numbers, with only a few exceptions, the separation bubble locations were predicted 

correctly as were the total pressure loss profiles, the pressure distributions on the airfoil, the transition 

locations, and the mean velocity profiles in the boundary layer. The Transition-SST model was 

developed using data from flows over less aggressive airfoils, so it was encouraging that it performed 

well predicting the flow over the L1A without any modification to the model. 

In the next phase of the work, flow control was added using vortex generator jets. Each blade in the 

cascade had a central cavity to which compressed air was supplied through high speed solenoid valves. A 

spanwise row of small holes was drilled into the suction side of each blade at the streamwise location 

corresponding to the inviscid pressure minimum. The holes had a compound angle of 90 degrees to the 

main flow and 30 degrees to the surface, and extended into the central cavity. When air was supplied to 

the cavity, vortex generator jets were produced from the small holes. By pulsing the supply air, the VGJs 

could be pulsed. Cases were documented with different combinations of VGJ amplitudes, pulsing 

frequencies, and duty cycles. Reynolds numbers from 25,000 to 100,000 were considered under both 

high and low free-stream turbulence conditions. Measurements were the same as in the baseline cases, 

and the velocity measurements were also ensemble averaged at different phases within the jet pulsing 

cycle. The VGJs successfully suppressed separation and reduced aerodynamic losses. General 

observations were that the vortex generator jets control separation if they are of high enough amplitude 

and pulsing frequency. Steady jets tend to delay separation, but the boundary layer does still separate at 

low Reynolds numbers. Blowing ratios (jet velocity/local freestream velocity) of 2 to 3 were needed to 

achieve any separation control at the lowest Reynolds number. Pulsed jets are more effective than steady 

jets, allowing control with lower velocity jets. The jet mass flow requirements are very low, and the 

required jet velocities are low enough to be practically feasible in most cases. With a dimensionless 

frequency (F=fLj-te/Uave where f is the pulsing frequency, Lj-te is the distance from the jets to the trailing 

edge, and Uave is the average freestream velocity) of 0.5 or higher with Re=25,000, separation could be 

controlled with a blowing ratio of 1. When Re=50,000, a dimensionless pulsing frequency of 0.3 was 

adequate with a blowing ratio of 1. At lower frequencies it was sometimes possible to achieve partial 

control, particularly if higher blowing ratios were used. Once separation was fully controlled, increasing 

the pulsing frequency or blowing ratio further had no additional benefit. Freestream turbulence had some 

effect in reducing separation, but the effect was small in comparison to the VGJ effect. 

Numerical simulations of cases with VGJs showed that unsteady RANS (URANS) calculations with 

the Transition-SST model failed to predict reattachment for cases in which the VGJs caused reattachment 

in the experiments. Although the Transition-SST model worked well without the VGJs, it failed to 

capture the effect of the jets. Large Eddy Simulations (LES) were then tried, and the results were 

generally better. Although the match to the experiments was not perfect, the LES generally predicted 

separation and reattachment in the correct cases to match the experiments. 

Cases with VGJs were followed by cases with unsteady wakes. A wake generator was constructed 

and installed in the wind tunnel. The wake generator had two chains which formed loops around the 

cascade near the top and bottom endwalls. Each chain traversed around a drive gear and six idler 

sprockets. The drive gears were on a common axle, so both chains moved together. The top and bottom 

chains were connected by carbon fiber rods which were spaced between 1 and 2 blade spacings apart. 

The rods moved across the inlet of the cascade, 0.54 axial chord upstream of the leading edges, and then 

returned to complete the loop downstream of the cascade. The rod shed wakes which simulated the 

wakes from an upstream blade row. Rod speed was set to produce flow coefficients ranging from 0.35 to 

1.4. Cases were documented with various flow coefficients and rod spacings at Re=25,000 and 50,000 

under both high and low background freestream turbulence conditions. A photo detector was used to 

NASA/CR—2012-217415 2
sense the passing of each wake, and the voltage signal from the photo detector was digitized along with 

hot wire data so that velocity results could be ensemble averaged at different phases within the wake 

passing events. The wakes caused the boundary layer to reattach, but if the time between wakes was 

sufficiently long, the boundary layer would separate. In agreement with the VGJ results, a dimensionless 

rod passing frequency, F (defined as described above for the pulsed VGJs), of about 0.5 was sufficient to 

keep the boundary layer attached over most of the airfoil surface when Re=25,000. For Re=50,000, a 

dimensionless wake passing frequency of 0.3 was sufficient. Whether the frequency was achieved by 

increasing the rod speed or by decreasing the rod spacing did not appear to matter. As with the VGJs, 

high background freestream turbulence had some effect in reducing separation, but the effect was small in 

comparison to the wake effect. The turbulence level in the rod wakes was about 14% at the leading edge 

of the cascade regardless of whether the turbulence level between wakes was low (0.6%) or high (4%). 

Cases with wakes were simulated using URANS and LES. Both were able to capture the wake effect 

and generally predicted separation and reattachment to occur in the right cases to match those experiments 

simulated. Quantitatively the results were mixed. In some cases the match between the calculations and 

experiments were good for pressure and velocity results, but in other cases there were significant differences. 

In the final phase of the study, wakes and VGJs were combined. The valves controlling the VGJs 

were triggered and timed using the signal from the photo detector used to sense the wake passing. 

Various timings were used for the VGJs with respect to the wakes, including cases in which the jet pulses 

coincided with the wakes and cases with the jets occurring between wakes. Cases with multiple jet pulses 

for each wake were also considered. Cases were acquired for Re=25,000 and 50,000 with both low and 

high background freestream turbulence. The timing of the jets with respect to the wakes had some effect, 

but in general once the disturbance frequency was high enough to control separation, the timing was not 

very important. Results were nearly identical for otherwise similar low and high freestream turbulence 

cases, indicating that the wake and VGJ effects overwhelmed the effect of the background turbulence. 

Contents of the Report 

Most of the results of the present study, including description of the experiments and computations, 

have been provided in a series of technical papers that were published during the course of the study. These 

papers were peer reviewed for various conferences and several have also become journal publications. 

These papers are included in this report as the primary documentation of the work. The final report is 

divided into three separate NASA Contractor Reports prepared under NASA Interagency Agreement 

No. NNC07IA10I (NASA/CR—2012-217415, NASA/CR—2012-217416, and NASA/CR—2012-217417). 

The following sections comprise Part I (NASA/CR—2012-217415) of this report: 

1. Brief description of experimental facility with pictures that did not appear in other publications. 

2. List of experimental cases. 

3. Lists of conference and journal papers. 

4. Appendix A—Reprints of Papers and Publications. 

a. Copies of conference papers. 

b. Copies of journal papers. 

5. Appendix B—Additional Information. 

a. Measurements of transient velocity during VGJ pulses measured with the main flow in 

the wind tunnel turned off. 

b. Discussion and figures for cases with VGJs and cases with wakes including data that did 

not all appear in other papers. Included are figures showing the effect of changing VGJ 

spanwise spacing. 

c. Phase averaged boundary layer velocity and turbulence profiles for cases with VGJs and 

wakes combined that did not all appear in other papers. 

d. Phase averaged velocity and turbulence contours from downstream of the cascade for 

cases with VGJs and wakes combined that did not all appear in other papers. 

NASA/CR—2012-217415 3
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NASA/CR—2012-217415 5
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NASA/CR—2012-217415 6
Wakes were generated with the mechanism described briefly above and in more detail in the papers 

below. Figure 1.4 shows views of the wake generator during assembly. The rods which produce the 

wakes are not present in the figure, but would extend vertically between the upper and lower chain. In 

Fig. 1.4c the traverse for probe motion is visible. It is located downstream of the cascade within the wake 

generator circuit. The traverse has three linear stages and a rotating stage. Details of the wakes produced 

by the rods, including the velocity deficit, turbulence level and spectra are provided in the papers below. 

Measurements were made near the midspan of the airfoils, where the flow was free of endwall 

effects. The center airfoil included static pressure taps which were used to document the pressure 

distribution on the blade. An upstream Pitot tube and a Kiel probe traversed downstream of the cascade 

were used to document the total pressure loss through the cascade. A hot wire probe was used to measure 

the velocity and turbulence at the cascade inlet, and was also traversed downstream of the cascade in the 

wakes of the airfoils and at six streamwise stations in the suction surface boundary layer. From the hot 

wire data, mean and fluctuating velocity and spectra were computed. 

2. Experimental Cases 

The following tables show the experimental cases documented and the types of data acquired for the 

cases. Baseline experiments were run with steady inflow to the cascade and without wakes or VGJs. For 

the cases with VGJs, B is the blowing ratio (jet velocity/local freestream velocity), f is the pulsing 

frequency, and D is the duty cycle. Those cases with y in the Cp column are cases where pressure 

distributions were acquired on the airfoil. The Loss column indicates cases in which the total pressure 

loss between the upstream Pitot tube and the traversing Kiel probe 0.6 axial chord downstream of the 

cascade was documented. The Downstream U, u′ column indicates that the hot wire probe was traversed 

across the cascade in the same position as the Kiel probe traverse. The B.L. Profiles column indicates 

that velocity profiles were acquired at six streamwise stations in the suction surface boundary layer. The 

stations extended from near the beginning of the adverse pressure gradient region to near the trailing 

edge. 

The cases with VGJs were done with a spanwise jet spacing of 10 hole diameters, which corresponds 

to 0.063 axial chord. A few cases were also considered with some of the holes covered, increasing the 

spacing to 20, 40 and 80 diameters. These cases were with Re=50,000 and low freestream turbulence. 

Steady blowing was done with a blowing ratio of 2, and pulsed blowing with a frequency of 24 Hz, duty 

cycle of 10% and blowing ratio of 1. For these cases, the Station 6 column indicates that velocity profiles 

were acquired at streamwise Station 6, near the trailing edge. Velocity profiles were not acquired at the 

other streamwise stations. Both the Station 6 velocity profiles and total pressure loss profiles were 

acquired at 5 different spanwise positions between two VGJ holes. The spanwise locations were labeled 

a-e, with location b being 1⁄4 hole pitch from location a, location c being 1⁄4 pitch from location b (and 1⁄2 

pitch from location a), location d being 1⁄4 pitch from location b and 3⁄4 pitch from location a, and location 

e being 1⁄4 pitch from d and one full pitch from location a. In general, there was not much variation across 

the span at the measurement locations considered. There was little difference between the results with 

10D and 20D spacing. With 40D spacing, a slightly higher blowing ratio was needed with the pulsed jets 

(1.5 vs. 1) to achieve the same separation control observed with the closer spacing. With 80D spacing, 

the jets still had some effect, but separation control was largely lost. These results indicate that the 10D 

spacing used in most of the experiments was more conservative than necessary. Unlike the rest of the 

results, which appear in the attached conference and journal papers, the results with different VGJ spacing 

do not appear in any papers. Some discussion and figures from these cases are in the last section below, 

and the data files from these cases are included with the other data files from the study. 

 

NASA/CR—2012-217415 7
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ivided by the 

 acquired at t

n ensemble av

and VGJs co

these cases a 

inlet velocity

.14. Without

and 50,000. N

s were chosen

e case with on

o pulses per w

as long as tho

ed to name the

he names was 

files included

on data files a

as shown abo

w Freestream

B 

0 

0 

0 

0 

0 

0 

0 

0 1 

0 2 

0 2.5 

0 3 

without VGJs, 

blade spacing

the inlet plane

veraged at dif

ombine were c

rod spacing o

y for a flow co

t VGJs, this w

Nine different

n to place the j

nly wakes. Ca

wake period. T

ose in (b-d). T

e cases in the

used when th

d with this rep

and in tables 

ove and used i

m Turbulence

f [Hz] D

 

 

 

 

 

 

 

 

 

 

 

 

 

the Spacing c

g in the casca

e of the cascad

fferent phases

considered las

of two blade s

oefficient of 

wake passing 

t VGJ timings

jet pulses at d

ases (b-d) hav

To maintain t

Timings (h-i) h

 papers subm

he data were o

port. The corr

below a

in papers a

e Baseline an

 [%] Cp

y 

y 

y 

y 

y 

y 

y 

y 

y 

y 

y 

column indic

ade. The colu

de at a fixed p

s during the w

st. Reynolds n

spacings was 

1.13. This ga

frequency wa

s relative to th

different time

ve a single pu

the same over

have three pu

mitted for publ

originally acq

responding na

b c d

c d e 

nd VGJ cases

Loss

y 

y 

y 

y 

y 

y 

y 

y 

y 

y 

ates the spaci

umn marked U

point between

wake passing c

numbers of 2

used. The ro

ave a dimensi

as too low for

he wakes wer

es within or be

lse for each w

rall blowing p

ulses per wake

lication and in

quired, and th

ames are as fo

e f g i 

f i b g 

s (no wakes) 

Downst

U, u' 

y 

y 

y 

y 

y 

y 

y 

ing between t

Upstream indi

n two blades.

cycle. 

5,000 and 50

od velocity wa

ionless wake 

r full separatio

re considered

etween wakes

wake passing.

period, the pu

e. 

ncluded below

hese original 

ollows. 

j 

h 

ream B.L. 

Profil

6 

statio

y 

y 

y 

y 

y 

y 

y 

y 

the 

icates 

 The 

,000 

as 

on 

, as 

s. 

ulses 

w. A 

es, 

ons 

NASA/CR—2012-217415 8
Case Re B f [Hz] D [%] Cp Loss Downstream 

U, u' 

B.L. 

Profiles, 

6 

stations 

Steady Jets 50000 0.25 y 

 50000 0.5 y 

 50000 0.75 y y 

 50000 1 y y 

 50000 1.5 y y 

 50000 2 y y y 

 

Steady Jets 100000 0.25 y y 

 100000 0.5 y y y 

 100000 0.75 y y y 

 100000 1 y y 

 

Pulsed Jets 25000 1 3 10 y 

 25000 2 3 10 y 

 25000 3 3 10 y y 

 

Pulsed Jets 25000 0.5 3 50 y 

 25000 0.75 3 50 y 

 25000 1 3 50 y y 

 25000 1.5 3 50 y 

 25000 2 3 50 y y 

 25000 2.5 3 50 y 

 25000 3 3 50 y y 

 

Pulsed Jets 50000 1 3 50 y 

Pulsed Jets 25000 0.5 6 10 y 

 25000 0.75 6 10 y 

 25000 1 6 10 y y y 

 25000 1.5 6 10 y y 

 25000 2 6 10 y y 

 25000 2.5 6 10 y 

 25000 3 6 10 y y 

 

Pulsed Jets 25000 0.5 6 50 y 

 25000 0.75 6 50 y 

 25000 1 6 50 y y 

 25000 1.5 6 50 y y 

 25000 2 6 50 y 

 25000 2.5 6 50 y 

 25000 3 6 50 y 

 

Pulsed Jets 50000 0.25 6 10 y 

 50000 0.5 6 10 y 

 50000 0.75 6 10 y 

NASA/CR—2012-217415 9
Case Re B f [Hz] D [%] Cp Loss Downstream 

U, u' 

B.L. 

Profiles, 

6 

stations 

 50000 1 6 10 y y y y 

 50000 1.5 6 10 y 

 50000 2 6 10 y y 

 

Pulsed Jets 50000 0.25 6 50 y 

 50000 0.5 6 50 y y 

 50000 0.75 6 50 y y 

 50000 1 6 50 y y y y 

 50000 1.5 6 50 y 

 50000 2 6 50 y 

 

Pulsed Jets 25000 0.25 12 10 y 

 25000 0.5 12 10 y y 

 25000 0.75 12 10 y y 

 25000 1 12 10 y y y 

 25000 1.5 12 10 y y 

 25000 2 12 10 y y 

 25000 2.5 12 10 y 

 25000 3 12 10 y y 

 

Pulsed Jets 25000 0.25 12 50 y 

 25000 0.5 12 50 y 

 25000 0.75 12 50 y 

 25000 1 12 50 y 

 25000 1.5 12 50 y y 

 25000 2 12 50 y 

 25000 2.5 12 50 y 

 25000 3 12 50 y 

 

Pulsed Jets 50000 0.25 12 10 y y 

 50000 0.5 12 10 y y y 

 50000 0.75 12 10 y y 

 50000 1 12 10 y y y 

 50000 1.5 12 10 y 

 50000 2 12 10 y y 

 

Pulsed Jets 50000 0.25 12 50 y y 

 50000 0.5 12 50 y y y 

 50000 0.75 12 50 y y 

 50000 1 12 50 y y 

 50000 1.5 12 50 y 

 50000 2 12 50 y y 

 50000 3 12 50 y 

 

NASA/CR—2012-217415 10
Case Re B f [Hz] D [%] Cp Loss Downstream 

U, u' 

B.L. 

Profiles, 

6 

stations 

Pulsed Jets 100000 0.25 12 10 y 

 100000 0.5 12 10 y 

 100000 0.75 12 10 y 

 100000 1 12 10 y 

 

Pulsed Jets 100000 0.25 12 50 y 

 100000 0.5 12 50 y 

 100000 0.75 12 50 y 

 100000 1 12 50 y 

 

Pulsed Jets 25000 0.25 24 10 y y 

 25000 0.5 24 10 y y 

 25000 0.75 24 10 y 

 25000 1 24 10 y y y 

 25000 1.5 24 10 y 

 25000 2 24 10 y y 

 25000 2.5 24 10 y 

 25000 3 24 10 y y 

 

Pulsed Jets 25000 0.25 24 50 y 

 25000 0.5 24 50 y 

 25000 0.75 24 50 y 

 25000 1 24 50 y 

 25000 1.5 24 50 y 

 25000 2 24 50 y 

 25000 2.5 24 50 y 

 25000 3 24 50 y 

 

Pulsed Jets 50000 0.25 24 10 y y 

 50000 0.5 24 10 y y y y 

 50000 0.75 24 10 y 

 50000 1 24 10 y y y y 

 50000 1.5 24 10 y 

 50000 2 24 10 y 

 

Pulsed Jets 50000 0.25 24 50 y y 

 50000 0.5 24 50 y 

 50000 0.75 24 50 y 

 50000 1 24 50 y y y 

 50000 2 24 50 y 

 

Pulsed Jets 100000 0.25 24 10 y 

 100000 0.5 24 10 y y 

 100000 0.75 24 10 y 

 100000 1 24 10 y 

NASA/CR—2012-217415 11
Case Re B f [Hz] D [%] Cp Loss Downstream 

U, u' 

B.L. 

Profiles, 

6 

stations 

 

Pulsed Jets 100000 0.25 24 50 y 

 100000 0.5 24 50 y 

 100000 0.75 24 50 y 

 100000 1 24 50 y 

 

Pulsed Jets 50000 1 48 10 y 

Pulsed Jets 100000 0.25 48 10 y 

 100000 0.5 48 10 y 

 100000 0.75 48 10 y 

 

Pulsed Jets 100000 0.25 48 50 y 

 100000 0.5 48 50 y 

 100000 0.75 48 50 y 

 100000 1 48 50 y 

High Freestream Turbulence Baseline and VGJ cases (no wakes) 

Case Re B f [Hz] D [%] Cp Loss Downstream 

U, u' 

B.L. 

Profiles, 

6 

stations 

Baseline 25000 y y y y 

 50000 y y y y 

 100000 y y y y 

 150000 y y y y 

 200000 y y y y 

 300000 y y y y 

 

Steady Jets 25000 1 y 

 25000 2 y 

 25000 2.5 y 

 25000 3 y y 

 

Steady Jets 50000 0.5 y 

 50000 0.75 y 

 50000 1 y y 

 50000 1.5 y y 

 50000 2 y y 

 

Pulsed Jets 25000 2 3 10 y y 

 25000 2.5 3 10 y 

 25000 3 3 10 y y 

 

NASA/CR—2012-217415 12
Case Re B f [Hz] D [%] Cp Loss Downstream 

U, u' 

B.L. 

Profiles, 

6 

stations 

Pulsed Jets 25000 1 3 50 y 

 25000 2 3 50 y y 

 25000 3 3 50 y y 

 

Pulsed Jets 25000 1 6 10 y 

 25000 1.5 6 10 y y 

 25000 2 6 10 y y y 

 25000 3 6 10 y y y 

 

Pulsed Jets 25000 1 6 50 y 

 25000 2 6 50 y y 

 25000 2.5 6 50 y 

 25000 3 6 50 y y 

 

Pulsed Jets 50000 0.25 6 10 y 

 50000 0.5 6 10 y 

 50000 0.75 6 10 y 

 50000 1 6 10 y y 

 50000 1.5 6 10 y y 

 50000 2 6 10 y y 

 

Pulsed Jets 50000 0.25 6 50 y 

 50000 0.5 6 50 y 

 50000 0.75 6 50 y 

 50000 1 6 50 y y 

 50000 1.5 6 50 y y 

 50000 2 6 50 y 

 

Pulsed Jets 25000 0.5 12 10 y 

 25000 0.75 12 10 y y 

 25000 1 12 10 y y y 

 25000 1.5 12 10 y 

 25000 2 12 10 y y y 

 25000 2.5 12 10 y 

 25000 3 12 10 y 

 

Pulsed Jets 25000 0.5 12 50 y 

 25000 0.75 12 50 y 

 25000 1 12 50 y y 

 25000 2 12 50 y y 

 

Pulsed Jets 50000 0.13 12 10 y 

 50000 0.25 12 10 y 

 50000 0.5 12 10 y 

 50000 0.75 12 10 y y 

NASA/CR—2012-217415 13
Case Re B f [Hz] D [%] Cp Loss Downstream 

U, u' 

B.L. 

Profiles, 

6 

stations 

 50000 1 12 10 y y y 

 50000 1.5 12 10 y 

 

Pulsed Jets 50000 0.25 12 50 y 

 50000 0.5 12 50 y y 

 50000 0.75 12 50 y y 

 50000 1 12 50 y y 

 

Pulsed Jets 25000 0.5 24 10 y 

 25000 0.75 24 10 y 

 25000 1 24 10 y y 

 25000 1.5 24 10 y 

 25000 2 24 10 y y 

 

Pulsed Jets 50000 0.25 24 10 y 

 0.5 24 10 y y 

 0.75 24 10 y y y 

 1 24 10 y y 

 

Pulsed Jets 50000 0.25 24 50 y 

 0.5 24 50 y y 

 0.75 24 50 y y 

Low Freestream Turbulence cases with different VGJ spacings (no wakes) 

Case VGJ 

spacing 

Re B f [Hz] D [%] Cp Loss Station 

6 

Pulsed Jet 1 50000 1 24 10 y y y 

Steady Jet 1 50000 2 y y y 

Pulsed Jet 2 50000 1 24 10 y y y 

Steady Jet 2 50000 2 y y y 

Pulsed Jet 4 50000 1 24 10 y y y 

Steady Jet 4 50000 2 y y y 

Pulsed Jet 4 50000 1.5 24 10 y y* y* 

Pulsed Jet 8 50000 2.3 24 10 y y* y 

 

*only one spanwise location 

 

NASA/CR—2012-217415 14
Cases with Wakes (no VGJs) 

Case Re Spacing/Pitc

h 

f [Hz] Upstrea

m U, u' 

Cp Loss Downstrea

m U, u' 

B.L. 

Profiles, 

6 

stations 

Low TI 25000 1.6 6 y y y y y 

 37500 1.6 9 y y y y y 

 50000 1.6 12 y y y y y 

 100000 1.6 12 y y y 

 200000 1.6 12 y y y 

 25000 1.6 12 y y 

 50000 1.6 6 y y y 

 200000 1.6 6 y y 

 

Low TI 25000 1 6 y y y y 

 37500 1 9 y y y 

 50000 1 12 y y y y 

 

Low TI 25000 2 6 y y y y y 

 37500 2 9 y y y y y 

 50000 2 12 y y y y y 

 

High TI 25000 1.6 6 y y y y 

 37500 1.6 9 y y y 

 50000 1.6 12 y y y y 

 50000 1.6 6 y y 

 

High TI 25000 1 6 y y y y 

 37500 1 9 y y y 

 50000 1 12 y y y y 

 

Cases with Wakes and VGJs 

Case Re Timing f [Hz] Cp Loss Downstream 

U, u' 

B.L. Profiles, 

6 stations 

Low TI Re=25000 a 3 y y y y 

 b 3 y y y y 

 c 3 y y y y 

 d 3 y y y y 

 e 3 y y y y 

 f 3 y y y y 

 i 3 y y y y 

 j 3 y y y y 

 50000 a 6 y y y y 

 b 6 y y y y 

 c 6 y y y y 

 d 6 y y y y 

 e 6 y y y y 

NASA/CR—2012-217415 15
 g 6 y y y y 

 

High TI Re=25000 a 3 y y y y 

 b 3 y y y y 

 c 3 y y y y 

 d 3 y y y y 

 e 3 y y y y 

 f 3 y y y y 

 50000 a 6 y y y y 

 b 6 y y y y 

 c 6 y y y y 

 d 6 y y y y 

 e 6 y y y y 

 g 6 y y 

 

NASA/CR—2012-217415 16
3. Lists of publications 

Dissertations and Theses 

1. Kartuzova, Olga V. “A Computational Study for the Utilization of Jet Pulsations in Gas Turbine 

Film Cooling and Flow Control,” PhD Dissertation, Cleveland State University, May 2010. 

2. Vinci, Samuel J. “CFD Simulations for the Effect of Unsteady Wakes on the Boundary Layer of a 

Highly Loaded Low Pressure Turbine Airfoil (L1A),” MS thesis, Cleveland State University, 

May 2011. 

Conference papers 

3. Volino, R.J. “Separated Flow Measurements on a Highly Loaded Low-Pressure Turbine Airfoil,” 

ASME paper GT2008-51445. Presented at the 2008 ASME International Gas Turbine 

Conference, Berlin, Germany. 

4. Ibrahim, M., Kartuzova, O., and Volino, R.J., “Experimental and Computational Investigations of 

Separation and Transition on a Highly Loaded Low-Pressure Turbine Airfoil: Part 1 – Low 

Freestream Turbulence Intensity,” ASME paper IMECE2008-68879. Presented at the ASME 

IMECE, Boston, MA, November 2008. 

5. Volino, R.J., Kartuzova, O., and Ibrahim, M., “Experimental and Computational Investigations of 

Separation and Transition on a Highly Loaded Low-Pressure Turbine Airfoil: Part 2 – High 

Freestream Turbulence Intensity,” ASME paper IMECE2008-68776. Presented at the ASME 

IMECE, Boston, MA, November 2008. 

6. Volino, R.J., Kartuzova, O., and Ibrahim, M.B., “Experimental and Computational Investigations 

of Low-Pressure Turbine Separation Control Using Vortex Generator Jets,” ASME paper 

GT2009-59983. Presented at the 2009 ASME International Gas Turbine Conference, Orlando, 

FL. 

7. Ibrahim, M.B., Kartuzova, O., and Volino, R.J., “LES and URANS Computational Investigations 

of LPT Blade (L1A) Separation Control using Vortex Generator Jets,” Proceedings of 

Turbulence, Heat and Mass Transfer 6, 2009. Presented at Turbulence, Heat and Mass Transfer 

6, Rome, Italy. 

8. Volino, R.J. and Ibrahim, M.B., “Separation Control on High Lift Low Pressure Turbine Airfoils 

Using Pulsed Jet Vortex Generator Jets,” Proceedings of the ASME-ATI-UIT 2010 Conference 

on Thermal and Environmental Issues in Energy Systems, 2010. Presented at ASME-ATI-UIT 

2010, Sorrento, Italy. 

9. Volino, R.J., Kartuzova, O., and Ibrahim, M.B., “Separation Control on a Very High Lift Low 

Pressure Turbine Airfoil Using Pulsed Vortex Generator Jets,” ASME paper GT2010-23567. 

Presented at the 2010 ASME International Gas Turbine Conference, Glasgow, UK. 

10. Volino, R.J., “Effect of Unsteady Wakes on Boundary Layer Separation on a Very High Lift Low 

Pressure Turbine Airfoil,” ASME paper GT2010-23573. Presented at the 2010 ASME 

International Gas Turbine Conference, Glasgow, UK. 

NASA/CR—2012-217415 17
11. Ibrahim, M., Kartuzova, O., Doucet, D., and Volino, R.J., “LES Flow Control Simulations for 

Highly Loaded Low Pressure Turbine Airfoil (L1A) Using Pulsed Vortex Generator Jets,” ASME 

paper GT2010-23015. Presented at the 2010 ASME International Gas Turbine Conference, 

Glasgow, UK. 

12. Volino, R.J., “Combined Effects of Wakes and Pulsed Vortex Generator Jet Flow Control on 

Boundary Layer Separation on a Very High Lift Low Pressure Turbine Airfoil,” ASME paper 

GT2011-46106. Presented at the 2011 ASME International Gas Turbine Conference, Vancouver, 

BC, Canada. 

13. Ibrahim, M., Vinci, S., Kartuzova, O., and Volino, R.J., “CFD Simulations of Unsteady Wakes on 

a Highly Loaded Low Pressure Turbine Airfoil (L1A),” ASME Paper GT2012-69770. Presented 

at the 2012 ASME International Gas Turbine Conference, Copenhagen. 

Other Conference Presentations 

14. Volino, R.J., Ibrahim, M.B. and Kartuzova, O. “Flow Control on Low-Pressure Turbine Airfoils 

Using Vortex Generator Jets,” In: 2009 Minnowbrook VI Workshop on Flow Physics and 

Control for Internal and External Aerodynamics, Blue Mountain Lake, NY, August 2009, 

NASA/CP-2010-216112 (Editors: Lagraff, J.A., Povinelli, L.A., Gostelow, J.P., and Glauser, M.). 

15. Volino, R.J. “Separation Control on a Cascade of Airfoils using Pulsed Vortex Generator Jets,” 

62nd Annual meeting of the Division of Fluid Dynamics of the American Physical Society, 

Minneapolis, MN, November 2009. 

Journal Papers 

16. Volino, R.J., “Separated Flow Measurements on a Highly Loaded Low-Pressure Turbine Airfoil,” 

ASME Journal of Turbomachinery, Vol. 132, paper 011007, 2010. 

17. Volino, R.J., Kartuzova, O., and Ibrahim, M.B., “Separation Control on a Very High Lift Low 

Pressure Turbine Airfoil Using Pulsed Vortex Generator Jets,” ASME Journal of 

Turbomachinery, Vol. 133, paper 041021-1, 2011. 

18. Volino, R.J., “Effect of Unsteady Wakes on Boundary Layer Separation on a Very High Lift Low 

Pressure Turbine Airfoil,” ASME Journal of Turbomachinery, Vol. 134, paper 011011-1, 2012. 

 

NASA/CR—2012-217415 18
Appendix A 

Reprints of Papers and Publications 

Copyrighted papers are reprinted with permission of the publisher. 

 

NASA/CR—2012-217415 19
1 Copyright © 2008 by ASME 

SEPARATED FLOW MEASUREMENTS ON A HIGHLY LOADED LOW-PRESSURE 

TURBINE AIRFOIL

Ralph J. Volino

Mechanical Engineering Department 

United States Naval Academy 

Annapolis, Maryland 21402 

volino@usna.edu 

ABSTRACT 

 Boundary layer separation, transition and reattachment have been 

studied on a new, very high lift, low-pressure turbine airfoil. 

Experiments were done under low freestream turbulence conditions on 

a linear cascade in a low speed wind tunnel. Pressure surveys on the 

airfoil surface and downstream total pressure loss surveys were 

documented. Velocity profiles were acquired in the suction side 

boundary layer at several streamwise locations using hot-wire 

anemometry. Cases were considered at Reynolds numbers (based on 

the suction surface length and the nominal exit velocity from the 

cascade) ranging from 25,000 to 330,000. In all cases the boundary 

layer separated, but at high Reynolds number the separation bubble 

remained very thin and quickly reattached after transition to 

turbulence. In the low Reynolds number cases, the boundary layer 

separated and did not reattach, even when transition occurred. This 

behavior contrasts with previous research on other airfoils, in which 

transition, if it occurred, always induced reattachment, regardless of 

Reynolds number. 

NOMENCLATURE 

Cf skin friction coefficient

Cp ( ) 2

T − ρU/PP2 e , pressure coefficient 

Cx axial chord length 

f frequency

H δ

∗ /θ, shape factor 

K ) /)(/( dsdUU 2 ν ∞∞ , acceleration parameter 

Ls suction surface length 

Lφ blade spacing (pitch) 

P pressure 

PS upstream static pressure 

PT upstream stagnation pressure 

PTe downstream stagnation pressure 

PSD power spectral density

Re LU se /ν , exit Reynolds number 

Reδ∗ displacement thickness Reynolds number 

Reθ momentum thickness Reynolds number 

s streamwise coordinate, distance from leading edge 

TI freestream turbulence intensity 

U mean streamwise velocity 

U∞ local freestream velocity 

Ue nominal exit freestream velocity, based on inviscid solution 

u′ time averaged rms streamwise fluctuating velocity 

x axial distance from leading edge 

y distance from wall 

Zw Zweifel coefficient 

α1 inlet flow angle 

α2 exit flow angle 

δ

∗ displacement thickness 

φ coordinate along blade spacing, normal to axial chord 

γ intermittency, fraction of time flow is turbulent 

ν kinematic viscosity 

ρ density 

θ momentum thickness 

ψ (PT-PTe)/(PT-PS), total pressure loss coefficient 

ψint total pressure loss integrated over blade spacing 

Subscripts 

p pressure minimum (suction peak) location 

pt distance from suction peak to transition start 

s separation location 

st distance from separation location to transition start

∞ free-stream 

INTRODUCTION 

 Boundary layer separation can lead to partial loss of lift and higher 

aerodynamic losses on low-pressure turbine (LPT) airfoils (e.g. 

Hourmouziadis [1], Mayle [2], and Sharma et al. [3]). As designers 

impose higher loading to improve efficiency and lower cost, the 

associated strong adverse pressure gradients on the suction side of the 

airfoil can exacerbate separation problems. The problem is 

particularly relevant in aircraft engines. Airfoils optimized for 

maximum power under takeoff conditions may still experience 

separation at cruise conditions, due to the lower density and therefore 

lower Reynolds numbers at altitude. A component efficiency drop of 

Proceedings of ASME Turbo Expo 2008: Power for Land, Sea and Air 

GT2008 

June 9-13, 2008, Berlin, Germany

GT2008-51445

This material is declared a work of the U.S. Government and is not subject to copyright protection in the United States.

Approved for public release; distribution is unlimited.

NASA/CR—2012-217415 20
2 Copyright © 2008 by ASME 

2% may occur between takeoff and cruise in large commercial 

transport engines, and the difference could be as large as 7% in smaller 

engines operating at higher altitudes [4, 5]. Prediction and control of 

suction side separation, without sacrifice of the benefits of higher 

loading, is therefore, crucial for improved engine design. 

 Separation and separated flow transition, which can lead to 

boundary layer reattachment, have received considerable attention. 

Studies have included flows over flat plates subject to pressure 

gradients similar to those on the suction side of LPT airfoils, and flows 

over airfoils either in single passage facilities or multi-blade cascades. 

Some have considered steady inflow conditions, while others have 

included the effect of unsteady wakes. Recent examples include [6, 7, 

8, 9, 10, 11, 12]. Volino [13] provides a review of some earlier 

studies. In general, previous work shows that the strong acceleration 

on the leading section of the airfoil keeps the boundary layer thin and 

laminar, even in the presence of elevated freestream turbulence. When 

separation does occur, it is usually just downstream of the suction 

peak. If transition then occurs in the shear layer over the separation 

bubble, it is typically rapid and causes the boundary layer to reattach 

[13, 14]. Transition is dependent on Reynolds number, freestream 

turbulence level, and the surface roughness conditions upstream of the 

separation point. Several correlations for separated flow transition 

have been developed based on experimental data, including those of 

Mayle [2], Hatman and Wang [15], Davis et al. [16], Yaras [17], 

Volino and Bohl [18], and Praisner and Clark [19]. 

 The advances in separation understanding and prediction have led to 

attempts at separation control. Zhang et al. [20], Bohl and Volino 

[21], Volino [22], and others provide examples using passive devices 

such as boundary layer trips. Others have used active devices such as 

vortex generator jets (e.g. Bons et al. [4], and Volino [23]) or plasma 

devices (e.g. Huang et al. [24]). 

 Another way to improve performance is to design airfoils with 

pressure gradients more resistant to separation, as described by 

Praisner and Clark [19]. Forward loading, for example, makes airfoils 

more separation resistant by extending the adverse pressure gradient 

on the aft portion of the suction side over a longer distance. This 

reduces the local pressure gradient at all locations, making separation 

less likely. If separation does occur, forward loading provides a longer 

distance along the airfoil surface for reattachment. Forward loading 

has some disadvantages, however. As noted by Zhang et al. [20], the 

longer region of turbulent flow on a forward loaded airfoil can lead to 

increased profile losses. Forward loading also creates longer regions 

of strong pressure gradient on the endwalls, which can produce 

stronger secondary flows and losses. If flow control were incorporated 

in the design of an advanced airfoil, as discussed by Bons et al. [25], it 

might be possible to produce an aft loaded airfoil that was resistant to 

separation and had low profile and secondary loss characteristics over 

a range of Reynolds numbers. 

 The objective of the present study is to document the flow over a 

very highly loaded LPT airfoil. The airfoil chosen was designed at the 

Air Force Research Laboratory (AFRL), and is designated the L1A. It 

is available on a limited basis to U.S. researchers from Clark [26]. 

Dimensions of the L1A as used in the present experiments are given in 

Table 1. Based on the design calculations of Clark [26], the L1A has 

10% higher loading than the “ultra-high lift” airfoils described by 

Zhang and Hodson [27], and 17% higher loading than the Pak B airfoil 

considered in several studies such as Volino [13, 14] and Bons et al. 

[4]. The design calculations indicate that the Zweifel coefficient 

increases from about 1.15 on the Pak B to about 1.35 on the L1A. If 

the definition for the Zweifel coefficient is taken as 

Zw=2 cos2

α2(Lφ/Cx)(tanα1+tanα2) (1) 

Table 1: Cascade parameters 

Axial 

Chord, Cx

[mm] 

True 

Chord 

[mm] 

Pitch, 

Lφ

[mm] 

Span 

[mm] 

Suction 

side, Ls

[mm] 

Inlet 

flow 

angle 

Exit 

flow 

angle 

134 146 136 724 203 35° 60° 

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

x/Cx

Cp

(a)

Pak B

L1M

L1A

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

(b)

Pak B

L1M

L1A

Fig. 1 Comparison of inviscid pressure profiles for the 

Pak B, L1M and L1A airfoils, a) Cp vs axial position, b) Cp

vs streamwise location 

as given by Lakshminarayana [28], Zw=1.08 for the Pak B (in 

agreement with the value given by McAuliffe and Sjolander [29]) and 

Zw=1.23 for the L1A. The L1A has the same inlet and exit flow 

angles and loading as the L1M airfoil used by Bons et al. [25]. The 

L1M is a mid-chord loaded design, and is resistant to separation even 

at very low Reynolds numbers. While the L1M and L1A were 

designed with the same methodology, the L1A is aft loaded, and is 

therefore expected to be more prone to separation. It is expected to be 

a good test airfoil for future flow control work, in which separation 

may be suppressed while maintaining the benefits of both very high 

loading and aft loading. Figure 1 shows the inviscid flow pressure 

profiles for the L1A, L1M and Pak B airfoils. The inviscid code used 

NASA/CR—2012-217415 21
3 Copyright © 2008 by ASME 

for the computations was adapted from Lewis [30]. Figure 1a shows 

the pressure coefficient Cp as a function of dimensionless axial 

position, x/Cx. These coordinates are useful for comparing the lift 

generated by the airfoils. Figure 1b shows Cp as a function of the 

streamwise distance along the airfoil surface, s/Ls, where Ls is the 

wetted surface length along the suction side. These coordinates are 

most useful for explaining the boundary layer development and are 

used to present the results below. The higher lift of the L1A and L1M 

compared to the Pak B is clear in Fig. 1. Also clear is the stronger 

adverse pressure gradient of the L1A on the downstream region of the 

suction side. The pressure gradient is shown again in Fig. 2 as the 

product of the local acceleration parameter K and the Reynolds 

number. The Reynolds number, Re, is based on Ls and the nominal 

exit velocity from the cascade Ue (computed using the inlet velocity 

and the design inlet and exit flow angles). The parameter K is 

inversely proportional to Re, so KRe is independent of Reynolds 

number. The negative K values downstream of s/Ls=0.7 are about 

twice as strong for the L1A than for the other two airfoils. This 

suggests more of a tendency for boundary layer separation and 

possibly less of a tendency for the flow to reattach. 

EXPERIMENTAL FACILITY AND MEASUREMENTS

 Experiments were conducted in a closed loop wind tunnel with a 

linear cascade in one corner of the loop. An axial fan sends air 

through turning vanes in the first corner of the tunnel and then through 

a heat exchanger. Following the heat exchanger is the second turn and 

a 6.25:1 area ratio contraction. Following the contraction is a 0.46 m 

square by 1.1 m long test section, which was not used in the present 

study. Following this test section is a 2.4 m long diffuser which 

expands the flow area to 0.75 m × 0.75 m. A fine screen with 0.12 

mm mesh thickness, 0.42 mm mesh spacing and 49% blockage is 

located in the exit plane of the diffuser. The screen breaks up the 

boundary layers which form on the diffuser walls and provides a 

uniform flow into the cascade test section. A seven blade cascade is 

located in the wind tunnel’s third turn, as shown in Fig. 3. A generic 

airfoil shape is shown in the figure. The flow continues from the 

cascade into the tunnel’s fourth turn, then through a second diffuser 

before returning to the fan. 

 The freestream turbulence entering the cascade was measured with a 

cross-wire probe positioned just upstream of the center blade. The 

streamwise velocity component had a turbulence intensity, TI, of 

0.8%, while the component intensity in the direction across the 

cascade was 0.5%. The spanwise component intensity was assumed to 

also be 0.5% based on the symmetry of the wind tunnel. The 

combined freestream turbulence intensity based on all three 

components was 0.6%. Spectral measurements showed that 80% of 

the energy in the streamwise fluctuations was due to unsteadiness at 

dimensionless frequencies fLs/Ue<2, where f is the frequency in Hz. 

Approximately 70% of the energy in the cross stream components was 

also below this frequency. If the turbulence is high pass filtered to 

remove this unsteadiness, the streamwise and cross stream component 

intensities are 0.36% and 0.27% respectively, for an overall freestream 

turbulence intensity of 0.3%. The integral length scale of the 

freestream turbulence is 6.3 cm in the streamwise direction and 6.7 cm 

in the other directions. While such low freestream turbulence and 

large length scales are not representative of engine conditions, they are 

still of interest. In zero or favorable pressure gradient boundary layers, 

high TI can cause bypass transition, but under adverse pressure 

gradients, natural transition appears to play a role at all TI levels. 

Hughes and Walker [31], for example observed evidence of TollmienSchlichting (TS) waves in cases with TI as high as 8%. Volino [14] 

also saw evidence of TS waves in both high and low TI cases. Low TI

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 −2

−1

0

1

2

3

4

5

s/Ls

KRe

Pak B

L1M

L1A

Fig. 2 Acceleration, KRe, vs streamwise location for Pak B, 

L1M and L1A airfoils based on inviscid solution 

tailboard

flap

flap

screen

B1

B2

B3

B4

B5

B6

B7

Fig. 3 Schematic of linear cascade 

cases provide a somewhat simpler environment for explaining the 

already complex phenomenon of separated flow transition. The 

present low TI cases will also serve as comparison cases for future 

work with elevated TI. Higher TI and length scales closer to the 

boundary layer thickness are expected to hasten the transition process. 

 The blades in the cascade were machined from high density foam, 

which has a consistency much like hard wood. The center blade, 

designated B4 in Fig. 3, contains pressure taps near the spanwise 

centerline. A tailboard, shown in Fig. 3, was needed to produce the 

correct exit flow angle from the cascade. Its position was set to 

produce periodicity at high Reynolds numbers. A tailboard on the 

opposite side of the cascade, and inlet guide vanes were found to be 

unnecessary. To produce the correct approach flow to the end blades 

NASA/CR—2012-217415 22
4 Copyright © 2008 by ASME 

(B1 and B7), the amount of flow escaping around the two ends of the 

cascade was controlled with the flaps shown in Fig. 3. The flap 

positions were set using a wool tuft upstream of each blade to check 

that the incoming flow approached the stagnation points with the 

correct angle. The inlet flow angle at the center of the cascade was 

also checked with a three-hole pressure probe and found to be within 

2° of the design angle. At high Reynolds numbers, the approach 

velocity to the middle four passages was measured to be uniform to 

within 6%, and the difference between any two adjacent passages was 

within 3%. At low Reynolds numbers, slightly more variation was 

observed, but the approach velocity to the middle two passages still 

agreed to within 5%. Good periodicity at high Reynolds numbers was 

also observed in the exit flow from the cascade, as evidenced by 

suction side velocity profiles acquired near the trailing edge of blades 

B2-B6, and by total pressure loss surveys, which are shown below. At 

low Reynolds numbers, when significant separation bubbles were 

present, the periodicity was not as good due to suppression of the 

separation bubble thickness on the blades closest to the tailboard. This 

is an unavoidable result when using a finite linear cascade to study 

separated flow. It is considered acceptable for the present facility, 

since its intended purpose is for the study of flow control, which if 

successful should suppress separation on all blades, thereby restoring 

periodicity even at low Reynolds numbers. 

Measurements 

 Pressure surveys were made using a pressure transducer (0-870 Pa 

range Validyne transducer). Stagnation pressure was measured with a 

pitot tube upstream of the cascade. Static pressure taps were located in 

the center blade as noted above. The uncertainty in the suction side 

pressure coefficients was 0.07. Most of this uncertainty was due to 

bias error. Stochastic error was minimized by averaging pressure 

transducer readings over a 10 second period. 

 A four component traverse with three linear stages and one rotating 

stage was located in the wind tunnel downstream of the cascade. The 

traverse produced an acceptably low blockage when it was located at 

least two axial chord lengths downstream of the cascade. Variations in 

the wind tunnel velocity were less than 2% as the traverse was moved 

to various positions. The traverse was used to hold and move probes 

for velocity and downstream pressure measurements. 

 Total pressure losses were documented using a Kiel probe traversed 

across three blade spacings, 0.63 Cx downstream of the cascade. To 

compute the mass flow averaged pressure loss, a velocity profile was 

also acquired along the same line as the total pressure surveys using a 

single sensor hot-film probe. 

 Velocity profiles on the suction surface were measured at the six 

streamwise stations listed in Table 2. All stations are downstream of 

the inviscid pressure minimum at s/Ls=0.49. Profiles were measured 

near the spanwise centerline of the airfoil. Profiles were acquired with 

a hot-wire anemometer (AA Lab Systems model AN-1003) and a 

single sensor hot-film probe (TSI model 1201-20). The sensor 

diameter is 51 μm, and the active length is 1.02 mm. At each 

measurement location, data were acquired for 26 seconds at a 20 kHz 

sampling rate (219 samples). All raw data were saved. The high 

sampling rate provides an essentially continuous signal, and the long 

sampling time results in low uncertainty in both statistical and spectral 

quantities. Data were acquired at 40 wall normal locations in each 

profile, extending from the wall to the free-stream, with most points 

concentrated in the near wall region. The probe was positioned as 

close to tangent to the airfoil surface as possible at each station, such 

that the probe body extended downstream of the sensor and the 

direction of the traverse was within 5° of normal to the surface. In 

most cases the closest point to the wall in each profile was within 

Table 2: Velocity profile measurement stations 

Station 1 2 3 4 5 6 

s/Ls 0.53 0.59 0.69 0.78 0.88 0.97 

x/Cx 0.65 0.72 0.80 0.86 0.92 0.97 

about 0.2 mm of the wall, which compares to boundary layer 

thicknesses ranging from 1.1 mm to over 40 mm. 

 Flow direction in a separation bubble cannot be determined with a 

single-sensor hot-wire, but velocity magnitude can be measured and 

was found to be near zero within the bubbles of the present cases when 

the flow was laminar. In cases where the flow became turbulent but 

remained separated, fluctuating velocities caused false high mean 

velocity readings in the separation bubble. With the exception of these 

turbulent separated cases, the uncertainty in the mean velocity is 3-5% 

except in the very near wall region, where near-wall corrections (Wills 

[32]) were applied to the mean velocity. Uncertainties in the 

momentum and displacement thicknesses computed from the mean 

profiles are 10%. Uncertainty in the shape factor, H, is 8%. Local 

skin friction coefficients were computed from the near wall mean 

velocity profiles using the technique of Volino and Simon [33]. This 

technique accounts for streamwise pressure gradient effects on the 

mean profile. The uncertainty in Cf is 8%. The uncertainty in the 

fluctuating streamwise velocity is below 10%. 

 Pressure surveys and velocity profiles were acquired at nominal 

Re=25,000, 50,000, 100,000, 200,000, 300,000, and 330,000. An 

additional pressure survey was acquired at Re=125,000. The Reynolds 

number, as defined above, is based on the suction surface length and 

the nominal cascade exit velocity. The corresponding Reynolds 

numbers based on the cascade inlet velocity and the axial chord length 

ranged from 10,000 to 133,000. 

RESULTS 

Pressure Profiles 

 Pressure profiles for all cases are shown in Fig. 4. The inviscid 

profile for the L1A airfoil is shown for comparison. At the three 

highest Re, with the exception of a slightly higher suction side peak, 

the data generally agree with the inviscid profile. This indicates that 

the boundary layer is attached over most of the airfoil in these cases. 

Some deviation between the three highest Re cases is visible at 

s/Ls=0.6 with the Cp value rising slightly at the lower Re. This may 

indicate a small separation bubble at this location. At the three lowest 

Reynolds numbers, the suction peak is suppressed and the Cp values 

have a constant value on the downstream part of the suction side. This 

plateau indicates that the boundary layer has separated. The Cp values 

never return to the inviscid line, indicating that the boundary layer 

never reattaches (i.e. the separation bubble bursts). At Reynolds 

numbers of 100,000 and below, the separation bubble always appeared 

to burst. At Re above 150,000 the boundary layer was attached over 

most of the airfoil. Between Re=100,000 and 150,000 the behavior 

was less predictable, with some dependence on whether the desired 

wind tunnel velocity was approached from above or below. The 

Re=124,500 case in Fig. 4 shows the suppressed suction peak and 

burst bubble of the lower Re cases, but in some trials (not shown) the 

Cp values dropped somewhat from their plateau near the trailing edge. 

 The pressure side Cp values also show some change with Reynolds 

number. Near the leading edge, Cp is higher for the lower Re cases, 

suggesting more of a leading edge separation bubble. Near the trailing 

edge the Cp values are again higher at the lower Re. This is believed 

to result when the suction side separation bubble forces fluid toward 

the pressure side, thereby increasing the pressure side velocity. 

NASA/CR—2012-217415 23
5 Copyright © 2008 by ASME 

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

Re=25,800

Re=50,900

Re=102,000

Re=124,500

Re=205,900

Re=297,300

Re=328,300

Inviscid

Fig. 4 Cp profiles 

0 0.5 1 1.5 2 2.5 3 3.5

x 105

0.9

1

1.1

1.2

1.3

Re

∫(Cpsuction−Cppressure)dx/Cx

Fig. 5 Lift based on integrated Cp profile 

 The lift on the airfoil can be determined by integrating the 

difference between the suction and pressure side Cp values along the 

axial direction. The result is shown as a function of Reynolds number 

in Fig. 5. The lift is about 20% lower for the separated flow cases. 

 The results in Fig. 4 contrast with results for the Pak B airfoil 

presented by Volino [13], Bons et al. [4], Simon et al. [34] and others. 

With low freestream turbulence, at Re≤50,000 the boundary layer on 

the Pak B airfoil separated and did not reattach, much like the L1A 

behavior of the present study. At Re≥300,000, the boundary layer 

remained attached over most of the surface, again similar to the L1A 

behavior shown above. At intermediate Reynolds numbers, however, 

the Pak B results showed separation with reattachment after a large 

separation bubble. At Re=100,000, for example, the boundary layer 

on the Pak B airfoil separated just after the suction peak and did not 

reattach until near the trailing edge. As Reynolds number increased, 

the reattachment location gradually moved upstream. The L1A does 

not show this large separation bubble behavior, instead appearing to 

switch more abruptly between a nearly fully attached boundary layer 

and a burst bubble. This will be explored in more detail with the 

velocity profiles below. 

Total Pressure Losses 

 The loss coefficient, ψ, is shown for all cases in Fig. 6. The 

coordinate φ indicates the distance in the direction perpendicular to the 

axial chord. The normalizing quantity Lφ is the blade spacing. The 

origin, φ=0, corresponds to the location directly downstream of the 

trailing edge of the center blade in the direction of the exit design flow 

−1.5 −1 −0.5 0 0.5 1 1.5 0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

φ/Lφ

ψ

B3

B4 B5

B6

Re=25,800

Re=50,900

Re=102,000

Re=124,500

Re=205,900

Re=297,300

Re=333,900

Fig. 6 Total pressure loss coefficient at 0.63 Cx

downstream of cascade 

0 0.5 1 1.5 2 2.5 3 3.5

x 105

0

0.1

0.2

0.3

0.4

0.5

Re

ψint

Fig. 7 Integrated total pressure loss coefficient for center 

blade as function of Re

angle. At the two highest Reynolds numbers the losses are low, and 

the location of the loss peaks are in the expected positions downstream 

of the airfoils. This indicates that the actual flow angle is essentially 

equal to the design angle. The peaks downstream of blades B3-B5 are 

similar to each other, indicating periodicity in the cascade. The loss 

coefficient is near zero between the peaks. At Re=205,900, the loss 

coefficients are slightly larger than at the higher Re. Although the 

difference is within the uncertainty, the trend is consistent with the 

appearance of a small separation bubble at this Reynolds number, seen 

in Fig. 4 at s/Ls=0.6. At the lower Reynolds numbers, the burst 

separation bubble results in much higher losses, and forces the peaks 

about 0.35Lφ toward the pressure side of each passage. This shift 

corresponds to a 30° change in the exit flow angle. The reduction in 

flow turning is consistent with the lower lift shown in Fig. 5. The 

peaks become noticeably smaller moving from B6 to B3, indicating 

the effect of the tailboard in reducing the separation bubble thickness. 

 The integrated loss around the center blade is computed as 

= ∫ ∫ − −

2/L

2/L

2/L

2/L

int Ud Ud

φ

φ

φ

φ

ψ φψ φ (2) 

and is shown in Fig. 7 as a function of Reynolds number. Losses 

increase with decreasing Re. Between Re=300,000 and 200,000, ψint

NASA/CR—2012-217415 24
6 Copyright © 2008 by ASME 

−1.5 −1 −0.5 0 0.5 1 1.5 0

0.05

0.1

0.15

0.2

0.25

φ/Lφ

u′/Ue

B3 B4 B5

B3

B4

B5

B6

Re=25,800

Re=50,900

Re=102,000

Re=124,500

Re=205,900

Re=297,300

Re=333,900

Fig. 8 Rms fluctuating streamwise velocity at 0.63 Cx

downstream of cascade 

increases by 37%, due to the thicker boundary layer at lower Reynolds 

numbers. The most dramatic increase, however, occurs between the 

attached and separated flow cases. The integrated loss increases by a 

factor of 7 between the highest and lowest Reynolds numbers. 

 The turbulence associated with the airfoil wakes is shown in Fig. 8. 

The rms fluctuating streamwise velocity at 0.63 Cx downstream of the 

trailing edge is normalized on the nominal exit velocity for each case. 

At the higher Reynolds numbers the boundary layer turbulence has 

decayed to about 4% of Ue. A double peak is apparent, with the 

slightly higher peak resulting from the suction side boundary layer and 

the smaller peak from the pressure side. Between blades the 

turbulence intensity is at the background level of the wind tunnel. At 

the lower Reynolds numbers, the turbulence level is much higher and 

in agreement with the loss coefficients, the peaks are shifted toward 

the pressure side of the passages. 

Velocity Profiles 

 Velocity profiles for the six suction surface measurement stations 

are shown in Fig. 9 for the nominal Re=25,000 case. The top row in 

the figure shows the distance from the wall normalized on the suction 

surface length plotted against the local mean velocity normalized on 

the nominal exit velocity, Ue. The boundary layer has just separated at 

the first measurement station and the separation bubble grows larger at 

the downstream stations. The boundary layer does not reattach. The 

second row in Fig. 9 shows the rms streamwise fluctuating velocity, u′, 

normalized with Ue. There is a very large peak located in the shear 

layer over the separation bubble which reaches a dimensionless 

magnitude of about 0.2. The third row in Fig. 9 shows the 

intermittency, γ. The intermittency is the fraction of time the flow is 

turbulent. It was determined at each measurement location based on 

the instantaneous streamwise velocity signal, using the technique 

described in Volino et al. [35]. Turbulent flow is defined here to 

include a range of large and small scale eddies. A boundary layer or 

shear layer may have significant u′ fluctuations but still be considered 

non-turbulent if these fluctuations are induced by an external source 

such as freestream turbulence or are associated with instability in a 

narrow frequency band. Transition to turbulence is characterized by 

the appearance of broadband fluctuations. In the intermittency 

processing routine, the velocity signal is high pass filtered, and the 

appearance of high frequency fluctuations is used to distinguish 

0 0 0 0 0 0 1 0

0.1

0.2

U/Ue

y/Ls

a) 1 2 3 4 5 6

0 0 0 0 0 0 0.1 0

0.1

0.2

u′/Ue

y/Ls

b) 1 2 3 4 5 6

0 0 0 0 0 0 1 0

0.1

0.2

γ

y/Ls

c) 1 2 3 4 5 6

Fig. 9 Profiles for Re=25,000 case: 

(a) mean velocity, (b) u′/Ue, (c) intermittency 

10−1 100 101 102

10−10

10−8

10−6

10−4

10−2

100

f⋅Ls/Ue

PSD/(Ls

⋅Ue

)

s/Ls=0.531

0.594

0.688

0.781

0.875

0.969

Fig. 10 Turbulence spectra for Re=25,000 case 

between turbulent and non-turbulent flow. The results in Fig. 9 

indicate that the shear layer remains non-turbulent until the last 

measurement station, where a small peak indicates the possible 

beginning of transition. 

 Figure 10 presents another view of the transition process using 

turbulence spectra. The spectra are computed from the fluctuating 

velocity signal acquired at the location of peak u′ at each measurement 

station shown in Fig. 9. Frequencies are resolved from 4.88 Hz to 10 

kHz in 4.88 Hz increments using 4096 point Fast Fourier Transforms 

to compute the spectra. The frequencies and power spectral density 

(PSD) are normalized using Ls and Ue. There is a clear broad band rise 

in the spectra between the third and fourth measurement stations, and 

the magnitude continues to rise downstream. The rise indicates that 

transition may be beginning by the fourth station. The rise in the 

spectra is not as abrupt as at higher Reynolds numbers (shown below), 

however, which may indicate that transition is only in its earliest 

stages. This may explain the intermittency values near zero in Fig. 9. 

 The velocity profiles for the Re=50,000 case are shown in Fig. 11. 

The mean and fluctuating velocity show essentially the same behavior 

observed in the Re=25,000 case of Fig. 9. The boundary layer does 

not reattach, in agreement with the pressure profile of Fig. 4. The 

intermittency values rise slightly above zero at the third station and 

indicate that transition is clearly underway by the fourth station. The 

NASA/CR—2012-217415 25
7 Copyright © 2008 by ASME 

0 0 0 0 0 0 1 0

0.1

0.2

U/Ue

y/Ls

a) 1 2 3 4 5 6

0 0 0 0 0 0 0.1 0

0.1

0.2

u′/Ue

y/Ls

b) 1 2 3 4 5 6

0 0 0 0 0 0 1 0

0.1

0.2

γ

y/Ls

c) 1 2 3 4 5 6

Fig. 11 Profiles for Re=50,000 case: 

(a) mean velocity, (b) u′/Ue, (c) intermittency 

10−1 100 101 102

10−10

10−8

10−6

10−4

10−2

100

f⋅Ls/Ue

PSD/(Ls

⋅Ue

)

s/Ls=0.531

0.594

0.688

0.781

0.875

0.969

Fig. 12 Turbulence spectra for Re=50,000 case 

corresponding turbulence spectra of Fig. 12 agree, showing a clear rise 

in the power spectrum between the second and third stations and a 

further rise to a more turbulent state by the fourth station. 

 The initiation of transition without boundary layer reattachment 

shown in Figs. 11 and 12 is markedly different than the behavior 

observed in previous studies with other LPT airfoils. Volino [13] 

noted that on the Pak B airfoil, when transition began in the separated 

shear layer, the associated mixing almost immediately induced 

boundary layer reattachment. This behavior was utilized by Volino 

[22] and by Zhang et al. [20], who used small passive devices on their 

airfoils to control separation. These devices were too small to trip the 

boundary layer to turbulent, but introduced small disturbances which 

accelerated the transition process in the shear layer, thereby moving 

reattachment upstream. Both Volino [22] and Zhang and Hodson [27] 

found that these small devices resulted in lower losses than large 

devices which immediately tripped the boundary layer to turbulent. 

The present results suggest that such devices may not work with the 

L1A airfoil, because transition is not sufficient to force reattachment. 

The strong negative acceleration parameter, shown in Fig. 2, 

particularly at low Re is apparently strong enough to prevent 

reattachment of the turbulent shear layer. 

 Figure 13 shows the velocity profiles for the Re=100,000 case. 

Similar to the previous cases, the separation bubble grows in the 

0 0 0 0 0 0 1 0

0.1

0.2

U/Ue

y/Ls

a) 1 2 3 4 5 6

0 0 0 0 0 0 0.1 0

0.1

0.2

u′/Ue

y/Ls

b) 1 2 3 4 5 6

0 0 0 0 0 0 1 0

0.1

0.2

γ

y/Ls

c) 1 2 3 4 5 6

Fig. 13 Profiles for Re=100,000 case: 

(a) mean velocity, (b) u′/Ue, (c) intermittency 

10−1 100 101 102

10−10

10−8

10−6

10−4

10−2

100

f⋅Ls/Ue

PSD/(Ls

⋅Ue

)

s/Ls=0.531

0.594

0.688

0.781

0.875

0.969

Fig. 14 Turbulence spectra for Re=100,000 case 

streamwise direction, the boundary layer does not reattach, and there is 

a very high peak in the fluctuating velocity in the shear layer over the 

separation bubble. The intermittency rises slightly above zero at the 

second station and indicates fully turbulent flow by the third station, 

where γ=1 in the shear layer. As in the previous case, the transition is 

not sufficient to induce boundary layer reattachment. Fig. 14 shows 

the spectra for the Re=100,000 case. There is a clear narrowband rise 

in the spectrum between the first and second stations, centered about a 

dimensionless frequency of 12. A further broadband rise occurs by the 

third station, indicating a turbulent shear layer, in agreement with the 

intermittency profiles of Fig. 13. The peak at the second station in 

Fig. 14 indicates a shear layer instability which likely initiates 

transition. Similar peaks are visible in the other Reynolds number 

cases, and will be discussed further below. 

 Velocity profiles for the Re=200,000 case are shown in Fig. 15. The 

boundary layer was laminar and on the verge of separation at the first 

two measurement stations. The small peak in intermittency at the first 

measurement station is believed to be erroneous, based on the spectra 

and the zero intermittency measurement at the second station. The 

skin friction, as determined from the near wall velocity profile, was 

near zero at the first two stations, but any separation bubble was very 

thin. Between the second and third stations the boundary layer 

underwent transition. The boundary layer was clearly attached and 

NASA/CR—2012-217415 26
8 Copyright © 2008 by ASME 

0 0 0 0 0 0 1 0

0.02

0.04

0.06

0.08

U/Ue

y/Ls

a) 1 2 3 4 5 6

0 0 0 0 0 0 0.1 0

0.02

0.04

0.06

0.08

u′/Ue

y/Ls

b) 1 2 3 4 5 6

0 0 0 0 0 0 1 0

0.02

0.04

0.06

0.08

γ

y/Ls

c) 1 2 3 4 5 6

Fig. 15 Profiles for Re=200,000 case: 

(a) mean velocity, (b) u′/Ue, (c) intermittency 

10−1 100 101 102

10−10

10−8

10−6

10−4

10−2

100

f⋅Ls/Ue

PSD/(Ls

⋅Ue

)

s/Ls=0.531

0.594

0.688

0.781

0.875

0.969

Fig. 16 Turbulence spectra for Re=200,000 case 

fully turbulent at the third station and remained attached at all 

downstream stations, in agreement with the pressure profiles of Fig. 4. 

The spectra for this case are shown in Fig. 16. The clear jump from 

laminar to turbulent flow is clear between the second and third 

stations. 

 The velocity profiles and spectra for the Re=300,000 case are shown 

in Figs. 17 and 18. The results are very similar to those of the 

Re=200,000 case, with transition occurring between the second and 

third stations, and the boundary layer remaining attached and turbulent 

downstream. 

 The shape factor, H, and skin friction coefficient Cf are good 

indicators of the state of the boundary layer with respect to separation 

and transition. The shape factor is shown for the high Re cases in Fig. 

19. It is between 2.5 and 3.5 at the first two measurement stations, 

where the boundary layer is laminar and on the verge of separation. It 

then drops rapidly with the onset of transition, reaching a minimum of 

about 1.57 at the fourth measurement station. While this value clearly 

indicates that the boundary layer is attached and turbulent, it is still 

above the zero pressure gradient turbulent value of about 1.4. At the 

downstream stations, H begins to rise, and it exceeds 1.7 by the last 

station. The skin friction coefficient is shown as a function of 

momentum thickness Reynolds number in Fig. 20. It rises from near 

zero after transition and decreases in the streamwise direction. Also 

0 0 0 0 0 0 1 0

0.02

0.04

0.06

0.08

U/Ue

y/Ls

a) 1 2 3 4 5 6

0 0 0 0 0 0 0.1 0

0.02

0.04

0.06

0.08

u′/Ue

y/Ls

b) 1 2 3 4 5 6

0 0 0 0 0 0 1 0

0.02

0.04

0.06

0.08

γ

y/Ls

c) 1 2 3 4 5 6

Fig. 17 Profiles for Re=300,000 case: 

(a) mean velocity, (b) u′/Ue, (c) intermittency 

10−1 100 101 102

10−10

10−8

10−6

10−4

10−2

100

f⋅Ls/Ue

PSD/(Ls

⋅Ue

)

s/Ls=0.531

0.594

0.688

0.781

0.875

0.969

Fig. 18 Turbulence spectra for Re=300,000 case 

0.5 0.6 0.7 0.8 0.9 1

1.5

2

2.5

3

3.5

s/Ls

H

Re=205,900

Re=297,300

Re=333,900

Fig. 19 Shape factor for high Re cases 

shown is a standard flat-plate correlation for turbulent boundary layers 

from Schlichting [36] and the Ludwieg-Tillmann correlation. The 

present Cf values are as much as 40% lower than the flat-plate 

correlation and agree well with the Ludweig-Tillmann correlation. 

The rising H and low Cf values are indicative of the strong adverse 

pressure gradient. The corresponding values on the Pak B airfoil [13] 

were closer to typical zero pressure gradient values, which is 

consistent with the weaker pressure gradient on the Pak B and the 

stronger tendency for reattachment at low Re on the Pak B. 

NASA/CR—2012-217415 27
9 Copyright © 2008 by ASME 

200 400 600 800 1000 1200 1400

0

2

4

6

8

x 10−3

Cf=0.0256/Reθ

0.25

Cf=0.246⋅ 10−0.678H Reθ

−0.268

Reθ

Cf

Re=205,900

Re=297,300

Re=333,900

Flat plate

Ludweig−Tillmann

Fig. 20 Skin friction coefficient for high Re cases 

Instability Prediction 

 The spectral peak noted above at the second station in Fig. 14 

suggests an instability in the shear layer that likely initiates transition. 

Although less distinct, similar peaks are visible just before or just after 

transition inception at all Reynolds numbers. Volino [14] observed 

similar peaks in the boundary layer on the Pak B airfoil and associated 

them with TS waves, using an analysis from Walker [37]. The TS 

waves are believed to form in the attached boundary layer between the 

beginning of the adverse pressure gradient region and the separation 

location. The TS waves then continued to grow in the separated shear 

layer until they became large enough to induce transition, as discussed 

in Volino and Bohl [18]. Walker [37] gave the most unstable 

frequency for TS waves as 

2/3* 2 Re2.3/2 − ∞ = δ πν Uf (3) 

Hughes and Walker [31] note that the frequency predicted by Eq. (3) is 

a function of streamwise position, since the freestream velocity, U∞, 

varies in non-zero pressure gradient cases, and the displacement 

thickness, δ

*

, changes as the boundary layer grows. Hence, a single 

frequency cannot be expected for all TS waves in a boundary layer. 

For the present study, U∞ and Reδ* were taken at the separation 

location. The Reδ* value was estimated with the boundary layer code 

TEXSTAN [38]. The code was used to a compute the laminar suction 

side boundary layer from the leading edge to the separation point using 

the pressure gradient from Fig. 4 as input for each case. The code 

stops when the flow separates, and the separation points computed 

were in good agreement with the pressure and velocity profiles 

presented above. Figure 21 shows the peak frequencies extracted from 

Figs. 10, 12, 14, 16 and 18 along with frequencies predicted by Eq. 

(3). The good agreement between the experimental and predicted 

values indicates that TS waves may play a role in the transition. It is 

also possible that other instability mechanisms are present. Stieger 

and Hodson [39], and Roberts and Yaras [40], for example, observed 

Kelvin-Helmholtz rollup of the shear layer. 

Transition Correlations 

 Several correlations for predicting the starting location for separated 

flow transition are available in the literature, and some of these are 

tested below against the present data. 

 Mayle [2] presents the following correlations for short and long 

separation bubbles. 

Rest = 300 Reθs

0.7 short bubble (4)

Rest=1000 Reθs

0.7 long bubble (5) 

200 300 400 500 600 700 800 900

0

500

1000

1500

2000

Reδ*

f [Hz]

Experiment

Walker [37] correlation

Fig. 21 Frequencies of spectral peaks along with predicted 

most unstable TS frequencies at separation location 

where Rest is the Reynolds number based on the freestream velocity at 

separation and the distance from separation to transition start. 

 The correlation of Davis et al. [16] is 

Rest = 25000 log10[coth(0.1732 TI)] (6) 

For the present experiments, TI could reasonably be set between 0.3% 

and 0.8%, as discussed above. Varying TI within this range does not 

significantly affect the correlation predictions. For Eq. (6) and all the 

correlations below, TI was set to 0.5%. 

 Hatman and Wang [15] identify several transition modes and 

present correlations for each of them. Their laminar separation mode 

transition correlation can be cast in terms of Rest as 

Rest = 0.0816 Res + 26805 (7) 

 Yaras [17] proposes 

Rest = 0.04 Res+6.3 × 104

[1 − tanh3(TF′)] (8) 

where TF′=max(TF, 1%), and TF=TI(ss/λ)

0.2, where ss is the distance 

from the leading edge to the separation point and λ is the integral 

length scale of the freestream turbulence. 

 Praisner and Clark [19] present the correlation 

Rest=173.0 Res Reθs

-1.227 (9) 

 The above correlations are based on the conditions at the separation 

location and in some cases the freestream turbulence intensity. Volino 

and Bohl [18] reasoned that instabilities begin to grow when the 

boundary layer becomes unstable at the start of the adverse pressure 

gradient region, and present the following correlation 

Rept=8.80 [6.37 – log10(TI2

)] Reθp

4/3 (10) 

where Rept is based on the freestream velocity at the suction peak and 

the distance from the suction peak to the transition start. 

 The location of the suction peak for each case was taken from the 

experimental data of Fig. 4. The separation location and the 

momentum thickness Reynolds numbers at separation, Reθs, and at the 

suction peak, Reθp, were taken from the TEXSTAN calculations 

described above. Equations (4) – (10) were then used to compute st, 

which is defined as the distance from the leading edge to the start of 

transition. The Reynolds numbers used with the correlations are listed 

in Table 3. The st results are shown as a function of Re in Fig. 22. 

The experimental transition locations, as determined from Figs. 9-18 

NASA/CR—2012-217415 28
10 Copyright © 2008 by ASME 

Table 3: Conditions at suction peak and separation location 

based on laminar boundary layer calculation 

Re sp/Ls ss/Ls Reθp Reθs

25,800 0.438 0.496 48 61 

50,900 0.438 0.494 68 85 

102,000 0.438 0.496 96 122 

205,900 0.493 0.538 151 182 

297,300 0.493 0.538 182 220 

333,900 0.493 0.537 193 231 

0 0.5 1 1.5 2 2.5 3 3.5

x 105

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

Re

s

t/Ls

Volino and Bohl [18]

Mayle [2] short bubble

Mayle [2] long bubble

Davis et al. [16]

Hatman and Wang [15]

Yaras [17]

Praisner and Clark [19]

Experiment

Fig. 22 Predicted and experimental transition start location, 

error bars indicate uncertainty due to finite spacing of 

measurement stations 

are also shown. The uncertainties in the experimental locations, which 

result from the finite spacing between the measurement stations, are 

indicated by the error bars in Fig. 22. Values of st/Ls>1 in the figure 

indicate that transition is predicted downstream of the trailing edge of 

the airfoil. 

 The Mayle [2] short bubble correlation, which is only intended for 

separation bubbles that reattach, predicts transition somewhat too far 

upstream at all Reynolds numbers. For the Re≥200,000 cases, all of 

the other correlations agree with the experimental results. For the 

lower Re cases, the Volino and Bohl [18] correlation continues to 

agree with the experiment, while the other correlations predict 

transition too far downstream. 

CONCLUSIONS 

 The flow over the very high lift L1A airfoil was studied 

experimentally using a linear cascade. Reynolds numbers based on 

suction surface length and nominal exit velocity ranging from 25,000 

to 330,000 were considered. In all cases the laminar suction surface 

boundary layer separated, but at Reynolds numbers greater than 

150,000 the separation bubble was very thin and short, and the 

boundary layer was attached over most of the surface. At lower 

Reynolds numbers the boundary layer separated and never reattached. 

Separation without reattachment caused the lift on the airfoil to 

decrease by 20% and the total pressure losses increased by a factor of 

7 above the highest Reynolds number case. Transition to turbulence 

occurred in all cases in the shear layer after separation, and appeared 

to be caused by the growth of Tollmien-Schlichting waves which 

originated in the boundary layer upstream of separation. The 

transition location was well predicted by the correlation of Volino and 

Bohl [18]. Transition caused immediate reattachment in the high 

Reynolds number cases, but the turbulent shear layer remained 

separated in the low Re cases. This behavior contrasts with previous 

studies on other LPT airfoils. In those studies, transition immediately 

triggered reattachment, even at low Reynolds numbers. On the present 

airfoil, the strong adverse pressure gradient prevents reattachment at 

low Re, even after transition occurs. 

ACKNOWLEDGMENTS

 This work was sponsored by the National Aeronautics and Space 

Administration. The grant monitor is Dr. Anthony Strasizar of the 

NASA Glenn Research Center. The support of the United States 

Naval Academy Technical Support Department Shop and Fluids 

Laboratory is greatly appreciated. 

 

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NASA/CR—2012-217415 29
11 Copyright © 2008 by ASME 

 

[14] Volino, R.J., 2002, “Separated Flow Transition Under Simulated 

Low-Pressure Turbine Airfoil Conditions: Part 2 –Turbulence 

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[22] Volino, R.J., 2003, “Passive Flow Control on Low-Pressure 

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of Turbomachinery, Wiley, New York. 

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[31] Hughes, J.D. and Walker, G.J., 2001, “Natural Transition 

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Turbomachinery, 123, pp. 392-401. 

[32] Wills, J.A.B., 1962, “The Correction of Hot-Wire Readings for 

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pp. 65-92. 

 

[33] Volino, R.J., and Simon, T.W., 1997, “Velocity and Temperature 

Profiles in Turbulent Boundary Layers Experiencing Streamwise 

Pressure Gradients,” ASME Journal of Heat Transfer, 119, pp. 

433-439. 

[34] Simon, T.W., Qiu, S., and Yuan, K., 2000, “Measurements in a 

Transitional Boundary Layer Under Low-Pressure Turbine 

Conditions,” NASA/CR-2000-209957. 

[35] Volino, R.J., Schultz, M.P., and Pratt, C.M., 2001, “Conditional 

Sampling in a Transitional Boundary Layer Under High FreeStream Turbulence Conditions,” ASME Journal of Fluids 

Engineering, 125, pp. 28-37. 

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[37] Walker, G.J., 1989, “Transitional Flow on Axial Turbomachine 

Blading,” AIAA Journal, 27, pp. 595-602. 

[38] Crawford, M.E., and Kays, W.M., 1976, “STAN5 – A Program 

for Numerical Computation of Two-Dimensional Internal and 

External Boundary Layer Flows,” NASA CR-2742. 

[39] Stieger, R.D., and Hodson, H.P., 2004, “The Transition 

Mechanism of Highly Loaded Low-Pressure Turbine Blades,” 

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Journal of Turbomachinery, 128, pp. 349-356. 

NASA/CR—2012-217415 30
1 Copyright © 2008 by ASME 

Proceedings of IMECE2008 

2008 ASME International Mechanical Engineering Congress and Exhibition 

November 2-6, 2008, Boston, Massachusetts, USA

 IMECE2008-68879 

EXPERIMENTAL AND COMPUTATIONAL INVESTIGATIONS OF SEPARATION AND 

TRANSITION ON A HIGHLY LOADED LOW-PRESSURE TURBINE AIRFOIL: PART 1 – 

LOW FREESTREAM TURBULENCE INTENSITY

Mounir Ibrahim and Olga Kartuzova 

Mechanical Engineering Department 

Cleveland State University 

Cleveland, Ohio 44115-2425 

(216) 687-2580, m.ibrahim@csuohio.edu 

Ralph J. Volino

Mechanical Engineering Department 

United States Naval Academy 

Annapolis, Maryland 21402-5042 

(410) 293-6520, volino@usna.edu 

ABSTRACT 

 Boundary layer separation, transition and reattachment have been 

studied on a new, very high lift, low-pressure turbine airfoil. 

Experiments were done under low freestream turbulence conditions on 

a linear cascade in a low speed wind tunnel. Pressure surveys on the 

airfoil surface and downstream total pressure loss surveys were 

documented. Velocity profiles were acquired in the suction side 

boundary layer at several streamwise locations using hot-wire 

anemometry. Cases were considered at Reynolds numbers (based on 

the suction surface length and the nominal exit velocity from the 

cascade) ranging from 25,000 to 330,000. In all cases the boundary 

layer separated, but at high Reynolds number the separation bubble 

remained very thin and quickly reattached after transition to 

turbulence. In the low Reynolds number cases, the boundary layer 

separated and did not reattach, even when transition occurred. 

 Three different CFD URANS (unsteady Reynolds averaged NavierStokes) models were utilized in this study (using Fluent CFD Code),

the k- shear stress transport model, the v2

-f k- model, and the 4 

equation Transition model of Menter. At Re=25,000, the Transition 

model seems to perform the best. At Re=100,000 the Transition model 

seems to perform the best also, although it under-predicts the pressure 

coefficient downstream of the suction peak. At Re= 300,000 all 

models perform very similar with each other. The Transition model 

showed a small bump in the pressure coefficient downstream from the 

suction peak indicating the presence of a small bubble at that location. 

NOMENCLATURE 

Cp   2

T  U/PP2 e , pressure coefficient 

Cx axial chord length 

Ls suction surface length 

L blade spacing (pitch) 

P pressure 

PS upstream static pressure 

PT upstream stagnation pressure 

PTe downstream stagnation pressure 

Re LU se / , exit Reynolds number 

Re U∞ / ν, momentum thickness Reynolds number 

s streamwise coordinate, distance from leading edge 

TI freestream turbulence intensity 

U mean streamwise velocity 

Ue nominal exit freestream velocity, based on inviscid solution 

U∞ freestream velocity 

u′ time averaged rms streamwise fluctuating velocity 

x axial distance from leading edge 

y distance from wall 

y+ distance from wall in viscous units 

1 inlet flow angle 

2 exit flow angle 

 coordinate along blade spacing, normal to axial chord 

 kinematic viscosity 

 density 

 momentum thickness 

 (PT-PTe)/(PT-PS), total pressure loss coefficient 

Subscripts 

p pressure minimum (suction peak) location 

pt distance from suction peak to transition start 

s separation location 

t transition start 

INTRODUCTION 

 Flow separation on the suction surface of low-pressure turbine 

(LPT) airfoils often occurs when turbine engines operate at low 

Reynolds numbers, as in the case of aircraft engines at high altitude 

cruise conditions. Low Reynolds numbers can cause the boundary 

layer to remain laminar and easily separate. This laminar separation 

results in an engine efficiency drop and an increase in fuel 

consumption (Mayle [1], Howell [2], and Singh [3]). 

 Simulation and prediction of transitional flow over LPT airfoils 

under a wide variety of Reynolds numbers, freestream turbulence 

parameters and with flow separation is essential for improvement in 

aircraft engine design. A great number of experimental and numerical 

investigations had been carried out in order to better understand the 

NASA/CR—2012-217415 31
2 Copyright © 2008 by ASME 

mechanisms of flow separation and transition on LPT airfoils. Volino 

[4] provides a review of some earlier studies. In general, previous 

work shows that the strong acceleration on the leading section of the 

airfoil keeps the boundary layer thin and laminar, even in the presence 

of elevated freestream turbulence. Some recent examples of those 

experimental studies are: Volino, [4,5], Mahallati et al.[6,7], Zoric et 

al. [8], and Zhang and Hodson [9]. Downstream of the suction peak 

the adverse pressure gradient can cause boundary layer separation, 

which may be followed by transition to turbulence and flow 

reattachment (Volino, [4,5]). 

 The advances in separation understanding and prediction have led to 

attempts at separation control. Zhang et al. [10], Bohl and Volino 

[11], Volino [12], and others provide examples using passive devices 

such as boundary layer trips. Others have used active devices such as 

vortex generator jets (e.g. Bons et al. [13], and Volino [14]) or plasma 

devices (e.g. Huang et al. [15]). 

 Another way to improve performance is to design airfoils with 

pressure gradients more resistant to separation, as described by 

Praisner and Clark [16]. Forward loading, for example, makes airfoils 

more separation resistant by extending the adverse pressure gradient 

on the aft portion of the suction side over a longer distance. This 

reduces the local pressure gradient at all locations, making separation 

less likely. If separation does occur, forward loading provides a longer 

distance along the airfoil surface for reattachment. Forward loading 

has some disadvantages, however. As noted by Zhang et al. [10], the 

longer region of turbulent flow on a forward loaded airfoil can lead to 

increased profile losses. Forward loading also creates longer regions 

of strong pressure gradient on the endwalls, which can produce 

stronger secondary flows and losses. If flow control were incorporated 

in the design of an advanced airfoil, as discussed by Bons et al. [17], it 

might be possible to produce an aft loaded airfoil that was resistant to 

separation and had low profile and secondary loss characteristics over 

a range of Reynolds numbers. 

 A reduction in the production costs as well as the weight of an 

engine can be achieved by increasing the loading on LPT airfoils, 

thereby allowing a reduction in the number of LPT blades. Therefore, 

very highly loaded airfoils are of great interest. Volino [18] 

experimentally studied the flow over a very highly loaded LPT airfoil, 

designed at the Air Force Research Laboratory (AFRL) and designated 

L1A. The L1A is available on a limited basis to US researchers from 

Clark [private communication]. Dimensions of the L1A as used in the 

present experiments are given in Table 1. 

Table 1: Cascade parameters for L1A airfoil 

Axial 

Chord, Cx

[mm] 

True 

Chord 

[mm] 

Pitch, 

L

[mm] 

Span 

[mm] 

Suction 

side, Ls

[mm] 

Inlet 

flow 

angle 

Exit 

flow 

angle 

134 146 136 724 203 35° 60° 

 Volino [18] reported that for the low Reynolds numbers considered 

(Re = 25,000 – 125000, based on exit velocity and suction side length) 

the flow separated and never reattached, even after transition to 

turbulence. For the higher Re (Re = 200,000 - 300,000) a very thin 

separation bubble was observed and the flow quickly reattached after 

transition occurred. Volino, [18] concluded that L1A differs from 

many previously studied LPT airfoils, where transition forced 

separated flow to reattach even at low Re. The L1A is considered to 

be a good airfoil for future flow control work, combining very high 

loading with a need for separation control. 

 Along with experimental work significant computational effort has 

been devoted to better understanding of separation and transition 

mechanisms in the LPT. McAuliffe and Yaras [19] conducted a Direct 

Numerical Simulation (DNS) study of laminar boundary layer 

separation on a flat surface in the presence of an adverse pressure 

gradient. Separation and transition mechanisms were examined under 

low and high disturbance environments. Singh [3] studied the flow 

physics in a LPT cascade under low Re number conditions using Large 

Eddy Simulation (LES). Calculations were carried out for Re = 10,000 

and 25,000 (based on inlet velocity and axial chord). The flow for both 

Reynolds numbers separated and never reattached. 

 DNS and LES calculation require high resolution grids, which 

results in high computational time, therefore these methods are very 

computationally expensive. This makes modification of Reynolds 

Averaged Navier-Stokes (RANS) methods to better predict separation 

and transition processes very attractive. Many studies have been done 

in the area of developing and testing transition models capable of 

accurate prediction of flow physics in turbomachinery. Some of the 

latest examples include Suzen et al. [20], who applied a transition 

model based on an intermittency transport equation to predictions of 

LPT experiments on the Pack B airfoil. Howell [2] used the 

Prescribed Unsteady Intermittency Model (PUIM) to study wake - 

surface flow interactions on a high lift LPT airfoil. This approach 

employs a set of correlations for transition onset and for spot 

production rate. A different approach was proposed by Menter et al. 

[21], based on two transport equations. The intermittency transport 

equation is used to trigger the transition onset. The transport equation 

for the transition momentum thickness Reynolds number (Ret) is used 

to capture non-local effect of freestream turbulence intensity and 

pressure gradient at the boundary layer edge. These two equations 

were coupled with a shear stress transport turbulence model (SST). 

This model was validated against experimental data for various 

turbomachinery and aerodynamic applications (Langtry et al. [22] and 

Menter et al. [21]). 

 The objective of the present work is to conduct a computational 

study of a very highly loaded low pressure turbine airfoil, designated 

the L1A under low freestream turbulence conditions. The 

experimental data have been presented previously (Volino [18]) and a 

summary is provided in this paper. On the computational side the 

problem has been extensively investigated utilizing different: 1) grid 

structures, 2) inlet velocity conditions, 3) turbulence models, and 4) 

Reynolds numbers. The present results will serve as baseline results 

for future work, in which attempts will be made to suppress separation 

through flow control at low Reynolds numbers. 

EXPERIMENTAL FACILITY AND MEASUREMENTS

 Experiments (described in details in Volino [18]) were conducted in 

a closed loop wind tunnel with a linear cascade in one corner of the 

loop. A seven blade cascade is located in the wind tunnel’s third turn, 

as shown in Fig. 1. A generic airfoil shape is shown in the figure. 

 The freestream turbulence entering the cascade was measured with a 

cross-wire probe positioned just upstream of the center blade. The 

streamwise velocity component had a turbulence intensity, TI, of 0.8% 

and integral length scale of 6.3 cm. These values were used in all the 

calculations described below. Further details of the freestream 

turbulence are in Volino [18]. While such low freestream turbulence 

and large length scales are not representative of engine conditions, 

they are still of interest as a limiting case. Also, in zero or favorable 

pressure gradient boundary layers, high TI can cause bypass transition, 

but under adverse pressure gradients, natural transition appears to play 

a role at all TI levels. 

 A tailboard, shown in Fig. 1, was needed to produce the correct exit 

flow angle from the cascade. Its position was set to produce 

periodicity at high Reynolds numbers. A tailboard on the opposite 

side of the cascade, and inlet guide vanes were found to be 

NASA/CR—2012-217415 32
3 Copyright © 2008 by ASME 

unnecessary. To produce the correct approach flow to the end blades 

(B1 and B7), the amount of flow escaping around the two ends of the 

cascade was controlled with the flaps shown in Fig. 1. The flap 

positions were set using a wool tuft upstream of each blade to check 

that the incoming flow approached the stagnation points with the 

correct angle. The inlet flow angle at the center of the cascade was 

also checked with a three-hole pressure probe and found to agree with 

the design angle to within about 2° of uncertainty. At high Reynolds 

numbers, the approach velocity to the middle four passages was 

measured to be uniform to within 6%, and the difference between any 

two adjacent passages was within 3%. At low Reynolds numbers, 

slightly more variation was observed, but the approach velocity to the 

middle two passages still agreed to within 5%. Good periodicity at 

high Reynolds numbers was also observed in the exit flow from the 

cascade, as evidenced by suction side velocity profiles acquired near 

the trailing edge of blades B2-B6, and by total pressure loss surveys, 

which are shown below. At low Reynolds numbers, when significant 

separation bubbles were present, the periodicity was not as good due 

to suppression of the separation bubble thickness on the blades closest 

to the tailboard. This deviation from periodicity is considered 

acceptable for the present facility, since its intended purpose is for the 

study of flow control, which if successful should suppress separation 

on all blades, thereby restoring periodicity even at low Reynolds 

numbers. 

tailboard

flap

flap

screen inlet

outlet

periodic

periodic

B1

B2

B3

B4

B5

B6

B7

Fig. 1 Schematic of a linear cascade of 7 airfoils (Volino 

[18]) with boundary conditions and computational domain 

used in present study, shown in dashed lines 

COMPUTATIONAL METHOD 

 The CFD predictions were performed with the numerical software 

tool FLUENT [23]. Predictions with three turbulence models were 

compared with the experimental results of Volino [18]. These models 

are described below. The computational domain consisted of two 

channels and the airfoil in the middle. It is shown on Fig. 1 in dashed 

lines. The boundary conditions applied along the sides of the domain 

are periodic, except along the airfoil surfaces, which are walls. Noslip boundary conditions were applied for all walls. 

 The inlet boundary condition was a prescribed uniform velocity, and 

is described in more detail below. The inlet into the two channels is 

located at 1.9 Cx upstream of the airfoil leading edge in the flow 

direction. At the outlet zero gage pressure was applied. The exit was 

located at 3.8 Cx downstream of the airfoil trailing edge in the flow 

direction. Different exit locations were tested to insure that 3.8 Cx was 

far enough downstream to achieve independent results (as indicated by 

pressure coefficients, pressure losses downstream of the cascade and 

velocity profiles on the airfoil) through the passages. 

 Unsteady calculations were performed for all cases. Convergence 

was established when: 1) residuals reduced to a value of 10-4, 2) 

velocity at the outlet and pressure on the airfoil suction side settled 

(fluctuations around certain mean values were observed, however 

there were no equal cycles), and 3) the mass imbalance was less than 

0.01 %. After convergence was achieved within each time step (t = 

0.0005 s) with the conditions listed above, time averaged results were 

obtained for 2000 time steps. 

Grid independence study 

 A grid independence study for 2D grids was conducted for Re = 

102,000. The V2F turbulence model (described below) was applied. 

Differences in grids are summarized in Table 2. Grids 2 and 3 showed 

maximum difference in Cp on the suction side of an airfoil less than 

2%, therefore grid 2 was used in this study. 

Table 2 Grid characteristics for grid independence study 

Grid 

number 

Size y+

 on the airfoil 

walls 

Number of 

points on 

airfoil ss 

Number of 

points on 

airfoil ps 

1 91516 2.95 

(range 0.1-8) 

290 240 

2 62469 0.20 

(range 0.01-0.6) 

290 240 

3 312393 0.20 

(range 0.01-0.6) 

1369 933 

Boundary conditions influence study 

 After grid independence was established several inlet velocity 

boundary conditions were tested. First, a uniform inlet velocity was 

applied in the direction of the design inlet flow angle (35°). This 

condition resulted in a slightly higher pressure on the leading section 

of the suction side of the airfoil compared to the experiment, 

indicating that the actual inlet angle could be different from the design 

angle. To investigate this possibility, a 2D inviscid calculation was 

done for the full cascade shown in Fig. 1, including the tailboard and 

flaps. The inlet velocity magnitude and direction were taken from the 

inviscid calculation along a line parallel to the blade row and 1.9 Cx

upstream of the blades in the flow direction, and used to set the inlet 

boundary conditions to the 2 channel domain described above. The 

inlet boundary conditions tested are summarized in Table 3, all were 

tested at Re=102,000. Four different inlet conditions are presented in 

Table 3. The first condition assumed a uniform inlet velocity at the 

design flow angle of 35°. The second condition used the velocity 

profile entering the two channel domain as obtained from the invisicid 

calculation. This condition shows a slight variation in the velocity 

profile at the inlet both in x and  directions with spatial averaged 

velocities Vx= 3.78 and V = 2.32 m/s. Accordingly, the spatial 

averaged inlet flow angle is 31.5°. The third condition used a uniform 

inlet velocity and flow angle (Vx= 3.78 and V = 2.32 m/s and inlet 

flow angle = 31.5°) based on the average values across the inlet of the 

two channel domain from the inviscid calculation. Thus the difference 

between conditions (2) and (3) are in the spatial variation in the inlet 

velocity while the averaged values are the same. The fourth condition 

NASA/CR—2012-217415 33
4 Copyright © 2008 by ASME 

used a uniform inlet velocity and angle based on the spatial averaged 

values across the full cascade (instead of averaging over two channels 

only as in condition (3)) from the inviscid calculation. Therefore the 

inlet velocity variations chosen in this exercise involve different inlet 

flow velocity magnitude and angle based on the design or the inviscid 

CFD results for the 7 blades (see Fig. 1). 

 Boundary condition 4 produced results in better agreement with the 

experimental data and was chosen for the rest of this investigation. 

The deviation of the inlet angle from the design angle is about 2°, 

which is within the uncertainty of the experimental measurement and 

therefore plausible. 

Table 3 Inlet boundary conditions for 2 channels tested, 

Re=102,000 

# Description Vx, 

m/s 

V, 

m/s 

Vmag, 

m/s 

1, 

deg 

1 Design condition 3.65 2.56 4.46 35 

2 Velocity profile from 

cascade simulation taken 

at inlet into 2 channels 

3.78 

(avg) 

2.32 

(avg) 

4.43 

(avg) 

31.5 

(avg) 

3 Velocity from cascade 

simulation averaged 

across inlet into 2 

channels 

3.78 2.32 4.43 31.5 

4 Velocity from cascade 

simulation averaged 

across full cascade at 

streamwise location of 

inlet into 2 channels. 

3.71 2.36 4.4 32.5 

Turbulence Modeling 

 In this study, the k  - sst model of Menter [24], the 

 fv 2 model of Durbin [25], and new Transition-sst (4 eq.) model 

of Menter [21], were compared for separated flow predictions on the 

highly loaded LPT airfoil. The unsteady Reynolds-averaged NavierStokes (URANS) equations were used as the transport equations for 

the mean flow. The governing equations and the detailed description 

of each model have been published elsewhere and therefore are not 

shown in this paper. 

 Shear-stress transport k  model (SKW-sst): This 

model, developed by Menter [24] is similar to the standard k  of 

Wilcox [26], but has an ability to account for the transport of the 

principal shear stress in adverse pressure gradient boundary layers. 

The model is based on Bradshaw's [27] assumption that the principal 

shear stress is proportional to the turbulent kinetic energy, which is 

introduced into the definition of the eddy-viscosity. These features 

make the SST k  model more accurate and reliable for a wider 

class of flows (e.g., adverse pressure gradient flows, airfoils, transonic 

shock waves) than the standard k  model [23]. 

 fv 2 model (V2F): According to Launder [28], the normal 

stress 

2

v , perpendicular to the local streamline plays the most 

important role to the eddy viscosity. Motivated by this idea, Durbin 

[25] devised a “four-equation” model, known as the 2   vk

model, or  fv 2 model. It eliminates the need to patch models in 

order to predict wall phenomena like heat transfer or flow separation. 

It makes use of the standard k   model, but extends it by 

incorporating the anisotropy of near-wall turbulence and non-local 

pressure strain effects, while retaining a linear eddy viscosity 

assumption. 

 Transition-sst (4 eq.) model: A new correlation-based 

transition model was proposed by Menter et al. [21]. This model is 

based on two transport equations. The intermittency transport 

equation is used to trigger the transition onset. The transport equation 

for the transition momentum thickness Reynolds number (Ret) is used 

to capture non-local effects of freestream turbulence intensity and 

pressure gradient at the boundary layer edge. Outside the boundary 

layer the transport variable was forced to follow the value of Ret

given by correlations. Those two equations were coupled with the 

shear stress transport turbulence model (SST). This model has recently 

become available in Fluent code.

RESULTS 

Re =25,000 

 Figure 2 shows the pressure coefficient plotted versus s/Ls along the 

suction and pressure surfaces for the turbulence models tested and Re 

= 25,000. Notice the abscissa is the distance along the blade surface 

normalized with respect to the suction surface length. The 

experimental data shows that the Cp values have a constant value on 

the downstream part of the suction side. This plateau indicates that the 

boundary layer has separated and that the boundary layer never 

reattaches (i.e. the separation bubble bursts); this of course refers to 

the time-averaged bubble as the bubble will probably burst 

sporadically. All turbulence models predict the pressure coefficient 

reasonably well with some deviation near the leading edge. The 

deviation could be partially attributed to uncertainty in the 

measurements and the inlet velocity profile as discussed above. All 

the models do well in predicting the separation location and the failure 

of the flow to reattach, as seen in the experiment. 

 Figure 3 shows the total pressure loss coefficient, , plotted versus 

dimensionless distance (/L) at a location 0.63 Cx downstream of the 

cascade. The definition of used in this study is the same as the one 

used in earlier work by Volino [18] and Bons et al. [29]. The 

coordinate  indicates the distance in the direction perpendicular to the 

axial chord. The normalizing quantity L is the blade spacing. The 

origin, =0, corresponds to the location directly downstream of the 

trailing edge of the center blade in the direction of the exit design flow 

angle. At low Reynolds numbers the burst separation bubble results in 

high losses and forces the peaks about 0.35L toward the pressure side 

of each passage. The peaks become noticeably smaller moving from 

B5 to B3, indicating the effect of the tailboard in reducing the 

separation bubble thickness. Due to the lack of periodicity in the 

experiment, the predicted loss coefficient is not expected to agree 

precisely with the experiment. The prediction is based on periodic 

boundary conditions and is not influenced by tailboard effects, so it 

should show higher losses and possibly lower flow turning (peaks 

shifted more to the left in the figure). This is indeed the case, as 

shown in Fig. 3. 

 Velocity profiles for the six suction surface measurement stations on 

blade B4 documented in the experiment are shown in Fig. 4 for the 

nominal Re=25,000 case. The figure shows the distance from the wall 

normalized on the suction surface length plotted against the local mean 

velocity normalized on the nominal exit velocity, Ue. The boundary 

layer has just separated at the first measurement station and the 

separation bubble grows larger at the downstream stations. 

NASA/CR—2012-217415 34
5 Copyright © 2008 by ASME 

0.00

0.50

1.00

1.50

2.00

2.50

0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0

s/Ls

Cp

Exp.

SKW-sst

Trans-sst

V2F

Fig. 2 Cp profiles, Re=25,000 

Fig. 3 Pressure loss coefficients at 0.63 Cx downstream of 

cascade, Re=25,000 

Fig. 4 Mean velocity profiles, Re = 25,000 

Fig. 5 Comparison of Transition-sst and experimental u′/Ue

profiles, Re= 25,000 

 

The boundary layer does not reattach. The velocity profiles at the six 

stations along the suction surface are predicted reasonably well by all 

models with the Transition model doing better overall. The prediction 

in the near wall region is different than the data since the 

measurements were done using hot-wire anemometry and therefore 

can not register negative velocities when separation occurs. 

 Figure 5 shows u′/Ue profile versus y/Ls at the six different stations 

along the suction side. u′ was obtained from CFD (using the Transition 

model only because of its best overall prediction) from the kinetic 

energy of turbulence and assuming an isotropic flow field (i.e. u′ = v′

=w′). Despite this assumption the CFD shows reasonable agreement 

with the data. The CFD shows a peak in the u′ profile that moves 

away from the wall as one travels from station (1) to (6). This peak 

value will be utilized (as will be shown later) to predict the start of 

transition. It is interesting that the experimental data show a very large 

peak located in the shear layer over the separation bubble which 

reaches a dimensionless magnitude of about 0.2. 

 At this Re (=25,000) the Transition model seems to perform the best 

(compared to SKW and V2F). 

Re =100,000 

 Figure 6 shows the pressure coefficient plotted versus s/Ls along the 

suction and pressure surfaces for the turbulence models tested and Re

= 100,000. All turbulence models predict the pressure coefficient very 

well except the Transition model, which shows under-prediction 

downstream from the suction peak. One possible explanation for this is 

that at s/Ls = 0.6 transition starts (as discussed below). The mixing 

associated with transition will tend to promote reattachment, which 

would result in a drop in the pressure coefficient. The simulation may 

be over predicting this tendency toward reattachment in this case, 

although the velocity profiles shown below do not indicate 

reattachment. 

NASA/CR—2012-217415 35
6 Copyright © 2008 by ASME 

0.00

0.50

1.00

1.50

2.00

2.50

0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0

s/Ls

Cp

Exp.

SKW-sst

Trans-sst

V2F

Fig. 6 Cp profiles, Re=100,000 

Fig. 7 Pressure loss coefficients at 0.63 Cx downstream of 

cascade, Re=100,000 

Fig. 8 Mean velocity profiles, Re = 100,000 

Fig. 9 Comparison of Transition-sst and experimental u′/Ue

profiles, Re= 100,000 

 Figure 7 shows the total pressure loss coefficient plotted versus 

dimensionless distance at 0.63 Cx downstream of the cascade. The loss 

coefficient was predicted reasonably well by all models with the 

Transition model showing the best agreement. As in the Re = 25,000 

case, the experimental results were not periodic, so precise agreement 

is not expected between the prediction and the data. 

 Figure 8 shows the velocity profiles (normalized with respect to the 

exit free stream velocity) versus y/Ls at the 6 stations along the suction 

side. The velocity profiles are predicted reasonably well by all models 

with the Transition model doing better overall. Again the near wall 

region is different than the data since the measurements were done 

using hot-wire anemometry and therefore can not register negative 

velocities when separation occurs. 

 Figure 9 shows u′/Ue profiles versus y/Ls at 6 stations along the 

suction side. As explained above, u′ was obtained from the CFD 

(using the Transition model) using the kinetic energy of turbulence 

and assuming an isotropic flow field. Again the CFD results show a 

peak in the u′ profile that moves away from the wall as one travels 

from station (1) to (6). Similarly, from the experimental data, the 

separation bubble grows in the streamwise direction, the boundary 

layer does not reattach, and there is a very high peak in the fluctuating 

velocity in the shear layer over the separation bubble. 

 At this Re (=100,000) the Transition model seems to perform the 

best (compared to SKW and V2F). Its one discrepancy is underprediction in the pressure coefficient downstream of the suction peak. 

Re =300,000 

 The experimental data for this case indicate that the boundary layer 

is attached over most of the airfoil. Figure 10 shows the pressure 

coefficient plotted versus s/Ls along the suction and pressure surfaces 

for the turbulence models tested and Re = 300,000. All turbulence 

models predict the pressure coefficient reasonably well including the 

area near the leading edge. The Transition model shows a bump at 

s/Ls = 0.6 indicating a small bubble that appears and then closes 

quickly. Although not as clear in the data, the predicted bubble may 

be correct. The predicted location is between two pressure 

measurement locations in the experiment, so the bubble presence 

would not necessarily be visible in the data. Also, a small bubble at 

this location was clearly noticed in the experimental data at a lower Re

(=200,000). 

NASA/CR—2012-217415 36
7 Copyright © 2008 by ASME 

0.00

0.50

1.00

1.50

2.00

2.50

0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0

s/Ls

Cp

Exp.

SKW-sst

Trans-sst

V2F 

Fig. 10 Cp profiles, Re=300,000 

Fig. 11 Pressure loss coefficients at 0.63 Cx downstream of 

cascade, Re=300,000 

Fig. 12 Mean velocity profiles, Re =300,000 

Fig. 13 Comparison of Transition-sst and experimental u′/Ue

profiles, Re= 300,000 

 Figure 11 shows the total pressure loss coefficient plotted versus 

dimensionless distance at 0.63 Cx downstream of the cascade. The 

loss coefficient was predicted reasonably well by all models. The 

experimental data showed periodic results at this higher Re. Therefore 

the periodic boundary condition applied in the CFD is consistent with 

the experiment. Nevertheless, the magnitude of the pressure loss 

coefficient is over predicted by about 15%. The location of the peaks 

is also shifted to the left of the experimental peaks in Fig. 11. The 

amount of the shift corresponds to about a 4° difference in flow angle. 

 Figure 12 shows the velocity profile (normalized with respect to the 

exit free stream velocity) versus y/Ls at 6 stations along the suction 

side. The velocity profiles are predicted very well by all models. 

Figure 13 shows u′/Ue profiles versus y/Ls at the 6 stations along the 

suction side. The CFD shows good agreement with the data. 

 At this Re (300,000) all models perform very similar with each 

other. The Transition model shows a small bump in the pressure 

coefficient downstream from the suction peak indicating the presence 

of a small bubble at that location. 

Prediction of Transition 

 From the above investigation one can conclude that the Transition 

model shows overall better agreement with the experimental data. 

Therefore in this section this model will be utilized to show how its 

predictions compare with data for the locations of separation and the 

start of transition. 

 Figure 14 shows contours of u′/Ue over the suction side of the airfoil 

overlapped with velocity vectors at 14a) Re=25,000, 14b) Re =100,000 

and 14c) Re = 300,000. On each plot the location of the 1) suction 

peak, 2) six stations used earlier in the velocity comparison with the 

data, 3) separation point and 4) transition start are shown. The CFD 

results of Fig. 14a show that the flow separates (at station (1)) with no 

reattachment, as observed experimentally. The location of transition 

was taken as the location where u′/Ue peaks in the shear layer over the 

separation bubble (between station (3) and (4) and close to (4)). In 

Fig. 14b the CFD data show that the flow separates (at station (1)) 

with no reattachment as observed experimentally. Again the location 

of transition was obtained where u′/Ue peaks in the shear layer 

(between station (1) and (2)). Notice that the suction peak is at about 

the same location as in the case with Re = 25,000. In Fig. 14c the CFD 

data show that with the possible exception of a small bubble appearing 

NASA/CR—2012-217415 37
8 Copyright © 2008 by ASME 

a) 

b) 

c) 

Fig. 14 Contours of u′/Ue, and velocity vectors (for 

Transition model) showing the location of: 1) suction peak, 

2) separation and 3) transition for a) Re = 25,000, b) Re = 

100,000, and c) Re = 300,000 

Fig. 15 Comparison between CFD data (Transition Model) 

and correlation for the start of transition 

at s/Ls = 0.6, the boundary layer is attached, which is consistent with 

the experimental data. Again the location of transition was obtained 

where the u′/Ue peaks (between station (2) and (3)). 

 To quantify the results in Fig. 14, Table 4 shows the numerical 

values of the predicted momentum thickness Reynolds number at the 

suction peak (Rep), the Reynolds number based on the freestream 

velocity at the suction peak and the streamwise distance from the 

suction peak to transition (Rept), and the streamwise locations of the 

suction peak, transition and separation. Table 5 shows corresponding 

measured quantities from the experiment. Note that Rep and ss in 

Table 5 were approximated using a laminar boundary layer calculation 

as explained in Volino [18]. The ranges given for Rept and st/Ls result 

from the finite spacing between measurement stations. The transition 

location is shown again in Fig. 15 along with a correlation from 

Volino and Bohl [30]: 

Rept=8.80 [6.37 – log10(TI2)] Rep

4/3

The agreement between the CFD and experiment shown in Tables 4 

and 5 and Fig. 15 is reasonably good. 

Table 4 CFD results for separation and transition locations 

(Transition Model) 

Re 25,000 100,000 300,000 

Rep 45 87 165 

Rept 11274 20816 73145 

sp/Ls 0.43 0.42 0.47 

st/Ls 0.74 0.59 0.64 

ss/Ls 0.53 0.53 0.6 

CONCLUSIONS 

 The flow over the very high lift L1A airfoil was studied 

experimentally using a linear cascade and reported by Volino [18]. 

Reynolds numbers based on suction surface length and nominal exit 

velocity ranging from 25,000 to 330,000 were considered. The 

experimental data showed that in all cases the laminar suction surface 

boundary layer separated, but at Reynolds numbers greater than 

150,000 the separation bubble was very thin and short, and the 

boundary layer was attached over most of the surface. At lower 

Transition 

Suction Peak Station 

1 

2 

3 

4 

5 

6 

Separation 

Suction Peak 

Transition 

Suction Peak Station 

1 

2 

3 

4 

5 

6 Separation 

Suction Peak 

Transition 

Station 

1 

2 

3 

4 

5 

6 

NASA/CR—2012-217415 38
9 Copyright © 2008 by ASME 

Reynolds numbers the boundary layer separated and never reattached. 

Transition to turbulence occurred in all cases in the shear layer after 

separation. Transition caused immediate reattachment in the high 

Reynolds number cases, but the turbulent shear layer remained 

separated in the low Re cases. 

 Three different CFD URANS models were utilized in this study, the 

SKW-sst, V2F and Transition models. At Re=25,000, the Transition 

model seems to perform the best. At Re=100,000 the Transition model 

again seems to perform the best, although it under-predicts in the 

pressure coefficient downstream of the suction peak. At Re= 300,000 

all models perform very similar with each other. The Transition 

model shows a small bump in the pressure coefficient downstream 

from the suction peak indicating the presence of a small bubble at that 

location. 

 Upon comparing the pressure loss coefficient at 0.63 Cx downstream 

of the cascade, the CFD shows a shift toward the pressure side of the 

passage compared to the data. Further investigation of the cause of 

this shift is needed. 

 Reasonably good agreement was obtained upon comparing the start 

of transition as obtained from CFD (using the Transition model), a 

published correlation and the experimental data. 

Table 5 Experimental Results for separation and transition 

locations 

Re 25,000 100,000 300,000 

Rep 48 96 193 

Rept 12140 

+/-3300 

28340 

+/-6500 

71170 

+/-22600 

sp/Ls 0.44 0.44 0.49 

st/Ls 0.78 

+/-0.094 

0.64 

+/-0.047 

0.64 

+/-0.047 

ss/Ls 0.50 0.50 0.54 

ACKNOWLEDGMENTS 

 This work was sponsored by the National Aeronautics and Space 

Administration. The grant monitor is Dr. Anthony Strasizar of the 

NASA Glenn Research Center. 

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Under a Wide Range of Operating Conditions,” AIAA Paper 

AIAA2003-3591. 

[21] Menter, F.R., Langtry, R.B., Likki, S.R., Suzen, Y.B., Huang, 

P.G., and Völker, S., 2006, “A Correlation based Transition 

Model using Local Variables Part 1- Model Formulation”, ASME 

Journal of Turbomachinery, 128, pp. 413-422. 

[22] Langtry, R.B., Menter, F.R., Likki, S.R., Suzen, Y.B., Huang, 

P.G., and Völker, S., 2006, “A Correlation based Transition 

Model using Local Variables Part II: Test Cases and Industrial 

Applications,” ASME Journal of Turbomachinery, 128, pp. 423-

444. 

[23] Fluent 6, 2001, Users manual, Fluent Inc., Lebanon NH, USA. 

[24] Menter, F. R., 1994, “Two-Equation Eddy-Viscosity Turbulence 

Models for Engineering Applications,” AIAA Journal, 32(8), pp. 

1598-1605. 

NASA/CR—2012-217415 39
10 Copyright © 2008 by ASME 

[25] Durbin, P.A., 1995, "Separated Flow Computations with the k--

v2 Model," AIAA Journal, 33(4), pp. 659-664. 

[26] Wilcox, D.C., 1998, Turbulent Modeling for CFD, DCW 

Industries, Inc., La Canada, California. 

[27] Bradshaw, P., Ferriss, D.H., and Atwell, N.P., 1967, “Calculation 

of Boundary-Layer Development Using the Turbulent Energy 

Equation,” Journal of Fluid Mechanics, 28(3), pp. 593-616. 

[28] Launder, B.E., and Spalding, D.B., 1974, “The Numerical 

Computation of Turbulent Flows,” Computer Methods in Applied 

Mechanics and Engineering, 3, pp. 269-289. 

[29] Bons, J.P., Pluim, J., Gompertz, K., Bloxham, M., and Clark, J.P., 

2008, “The Application of Flow Control to an Aft-Loaded Low 

Pressure Turbine Cascade with Unsteady Wakes,” ASME Paper 

GT2008-50864. 

[30] Volino, R.J., and Bohl, D.G., 2004, “Separated Flow Transition 

Mechanisms and Prediction With High and Low Freestream 

Turbulence Under Low Pressure Turbine Conditions,” ASME 

Paper GT2004-63360. 

 

NASA/CR—2012-217415 40
1 Copyright © 2008 by ASME 

Proceedings of IMECE2008 

2008 ASME International Mechanical Engineering Congress and Exhibition 

November 2-6, 2008, Boston, Massachusetts, USA

 IMECE2008-68879 

EXPERIMENTAL AND COMPUTATIONAL INVESTIGATIONS OF SEPARATION AND 

TRANSITION ON A HIGHLY LOADED LOW-PRESSURE TURBINE AIRFOIL: PART 1 – 

LOW FREESTREAM TURBULENCE INTENSITY

Mounir Ibrahim and Olga Kartuzova 

Mechanical Engineering Department 

Cleveland State University 

Cleveland, Ohio 44115-2425 

(216) 687-2580, m.ibrahim@csuohio.edu 

Ralph J. Volino

Mechanical Engineering Department 

United States Naval Academy 

Annapolis, Maryland 21402-5042 

(410) 293-6520, volino@usna.edu 

ABSTRACT 

 Boundary layer separation, transition and reattachment have been 

studied on a new, very high lift, low-pressure turbine airfoil. 

Experiments were done under low freestream turbulence conditions on 

a linear cascade in a low speed wind tunnel. Pressure surveys on the 

airfoil surface and downstream total pressure loss surveys were 

documented. Velocity profiles were acquired in the suction side 

boundary layer at several streamwise locations using hot-wire 

anemometry. Cases were considered at Reynolds numbers (based on 

the suction surface length and the nominal exit velocity from the 

cascade) ranging from 25,000 to 330,000. In all cases the boundary 

layer separated, but at high Reynolds number the separation bubble 

remained very thin and quickly reattached after transition to 

turbulence. In the low Reynolds number cases, the boundary layer 

separated and did not reattach, even when transition occurred. 

 Three different CFD URANS (unsteady Reynolds averaged NavierStokes) models were utilized in this study (using Fluent CFD Code),

the k- shear stress transport model, the v2

-f k- model, and the 4 

equation Transition model of Menter. At Re=25,000, the Transition 

model seems to perform the best. At Re=100,000 the Transition model 

seems to perform the best also, although it under-predicts the pressure 

coefficient downstream of the suction peak. At Re= 300,000 all 

models perform very similar with each other. The Transition model 

showed a small bump in the pressure coefficient downstream from the 

suction peak indicating the presence of a small bubble at that location. 

NOMENCLATURE 

Cp   2

T  U/PP2 e , pressure coefficient 

Cx axial chord length 

Ls suction surface length 

L blade spacing (pitch) 

P pressure 

PS upstream static pressure 

PT upstream stagnation pressure 

PTe downstream stagnation pressure 

Re LU se / , exit Reynolds number 

Re U∞ / ν, momentum thickness Reynolds number 

s streamwise coordinate, distance from leading edge 

TI freestream turbulence intensity 

U mean streamwise velocity 

Ue nominal exit freestream velocity, based on inviscid solution 

U∞ freestream velocity 

u′ time averaged rms streamwise fluctuating velocity 

x axial distance from leading edge 

y distance from wall 

y+ distance from wall in viscous units 

1 inlet flow angle 

2 exit flow angle 

 coordinate along blade spacing, normal to axial chord 

 kinematic viscosity 

 density 

 momentum thickness 

 (PT-PTe)/(PT-PS), total pressure loss coefficient 

Subscripts 

p pressure minimum (suction peak) location 

pt distance from suction peak to transition start 

s separation location 

t transition start 

INTRODUCTION 

 Flow separation on the suction surface of low-pressure turbine 

(LPT) airfoils often occurs when turbine engines operate at low 

Reynolds numbers, as in the case of aircraft engines at high altitude 

cruise conditions. Low Reynolds numbers can cause the boundary 

layer to remain laminar and easily separate. This laminar separation 

results in an engine efficiency drop and an increase in fuel 

consumption (Mayle [1], Howell [2], and Singh [3]). 

 Simulation and prediction of transitional flow over LPT airfoils 

under a wide variety of Reynolds numbers, freestream turbulence 

parameters and with flow separation is essential for improvement in 

aircraft engine design. A great number of experimental and numerical 

investigations had been carried out in order to better understand the 

NASA/CR—2012-217415 41
2 Copyright © 2008 by ASME 

mechanisms of flow separation and transition on LPT airfoils. Volino 

[4] provides a review of some earlier studies. In general, previous 

work shows that the strong acceleration on the leading section of the 

airfoil keeps the boundary layer thin and laminar, even in the presence 

of elevated freestream turbulence. Some recent examples of those 

experimental studies are: Volino, [4,5], Mahallati et al.[6,7], Zoric et 

al. [8], and Zhang and Hodson [9]. Downstream of the suction peak 

the adverse pressure gradient can cause boundary layer separation, 

which may be followed by transition to turbulence and flow 

reattachment (Volino, [4,5]). 

 The advances in separation understanding and prediction have led to 

attempts at separation control. Zhang et al. [10], Bohl and Volino 

[11], Volino [12], and others provide examples using passive devices 

such as boundary layer trips. Others have used active devices such as 

vortex generator jets (e.g. Bons et al. [13], and Volino [14]) or plasma 

devices (e.g. Huang et al. [15]). 

 Another way to improve performance is to design airfoils with 

pressure gradients more resistant to separation, as described by 

Praisner and Clark [16]. Forward loading, for example, makes airfoils 

more separation resistant by extending the adverse pressure gradient 

on the aft portion of the suction side over a longer distance. This 

reduces the local pressure gradient at all locations, making separation 

less likely. If separation does occur, forward loading provides a longer 

distance along the airfoil surface for reattachment. Forward loading 

has some disadvantages, however. As noted by Zhang et al. [10], the 

longer region of turbulent flow on a forward loaded airfoil can lead to 

increased profile losses. Forward loading also creates longer regions 

of strong pressure gradient on the endwalls, which can produce 

stronger secondary flows and losses. If flow control were incorporated 

in the design of an advanced airfoil, as discussed by Bons et al. [17], it 

might be possible to produce an aft loaded airfoil that was resistant to 

separation and had low profile and secondary loss characteristics over 

a range of Reynolds numbers. 

 A reduction in the production costs as well as the weight of an 

engine can be achieved by increasing the loading on LPT airfoils, 

thereby allowing a reduction in the number of LPT blades. Therefore, 

very highly loaded airfoils are of great interest. Volino [18] 

experimentally studied the flow over a very highly loaded LPT airfoil, 

designed at the Air Force Research Laboratory (AFRL) and designated 

L1A. The L1A is available on a limited basis to US researchers from 

Clark [private communication]. Dimensions of the L1A as used in the 

present experiments are given in Table 1. 

Table 1: Cascade parameters for L1A airfoil 

Axial 

Chord, Cx

[mm] 

True 

Chord 

[mm] 

Pitch, 

L

[mm] 

Span 

[mm] 

Suction 

side, Ls

[mm] 

Inlet 

flow 

angle 

Exit 

flow 

angle 

134 146 136 724 203 35° 60° 

 Volino [18] reported that for the low Reynolds numbers considered 

(Re = 25,000 – 125000, based on exit velocity and suction side length) 

the flow separated and never reattached, even after transition to 

turbulence. For the higher Re (Re = 200,000 - 300,000) a very thin 

separation bubble was observed and the flow quickly reattached after 

transition occurred. Volino, [18] concluded that L1A differs from 

many previously studied LPT airfoils, where transition forced 

separated flow to reattach even at low Re. The L1A is considered to 

be a good airfoil for future flow control work, combining very high 

loading with a need for separation control. 

 Along with experimental work significant computational effort has 

been devoted to better understanding of separation and transition 

mechanisms in the LPT. McAuliffe and Yaras [19] conducted a Direct 

Numerical Simulation (DNS) study of laminar boundary layer 

separation on a flat surface in the presence of an adverse pressure 

gradient. Separation and transition mechanisms were examined under 

low and high disturbance environments. Singh [3] studied the flow 

physics in a LPT cascade under low Re number conditions using Large 

Eddy Simulation (LES). Calculations were carried out for Re = 10,000 

and 25,000 (based on inlet velocity and axial chord). The flow for both 

Reynolds numbers separated and never reattached. 

 DNS and LES calculation require high resolution grids, which 

results in high computational time, therefore these methods are very 

computationally expensive. This makes modification of Reynolds 

Averaged Navier-Stokes (RANS) methods to better predict separation 

and transition processes very attractive. Many studies have been done 

in the area of developing and testing transition models capable of 

accurate prediction of flow physics in turbomachinery. Some of the 

latest examples include Suzen et al. [20], who applied a transition 

model based on an intermittency transport equation to predictions of 

LPT experiments on the Pack B airfoil. Howell [2] used the 

Prescribed Unsteady Intermittency Model (PUIM) to study wake - 

surface flow interactions on a high lift LPT airfoil. This approach 

employs a set of correlations for transition onset and for spot 

production rate. A different approach was proposed by Menter et al. 

[21], based on two transport equations. The intermittency transport 

equation is used to trigger the transition onset. The transport equation 

for the transition momentum thickness Reynolds number (Ret) is used 

to capture non-local effect of freestream turbulence intensity and 

pressure gradient at the boundary layer edge. These two equations 

were coupled with a shear stress transport turbulence model (SST). 

This model was validated against experimental data for various 

turbomachinery and aerodynamic applications (Langtry et al. [22] and 

Menter et al. [21]). 

 The objective of the present work is to conduct a computational 

study of a very highly loaded low pressure turbine airfoil, designated 

the L1A under low freestream turbulence conditions. The 

experimental data have been presented previously (Volino [18]) and a 

summary is provided in this paper. On the computational side the 

problem has been extensively investigated utilizing different: 1) grid 

structures, 2) inlet velocity conditions, 3) turbulence models, and 4) 

Reynolds numbers. The present results will serve as baseline results 

for future work, in which attempts will be made to suppress separation 

through flow control at low Reynolds numbers. 

EXPERIMENTAL FACILITY AND MEASUREMENTS

 Experiments (described in details in Volino [18]) were conducted in 

a closed loop wind tunnel with a linear cascade in one corner of the 

loop. A seven blade cascade is located in the wind tunnel’s third turn, 

as shown in Fig. 1. A generic airfoil shape is shown in the figure. 

 The freestream turbulence entering the cascade was measured with a 

cross-wire probe positioned just upstream of the center blade. The 

streamwise velocity component had a turbulence intensity, TI, of 0.8% 

and integral length scale of 6.3 cm. These values were used in all the 

calculations described below. Further details of the freestream 

turbulence are in Volino [18]. While such low freestream turbulence 

and large length scales are not representative of engine conditions, 

they are still of interest as a limiting case. Also, in zero or favorable 

pressure gradient boundary layers, high TI can cause bypass transition, 

but under adverse pressure gradients, natural transition appears to play 

a role at all TI levels. 

 A tailboard, shown in Fig. 1, was needed to produce the correct exit 

flow angle from the cascade. Its position was set to produce 

periodicity at high Reynolds numbers. A tailboard on the opposite 

side of the cascade, and inlet guide vanes were found to be 

NASA/CR—2012-217415 42
3 Copyright © 2008 by ASME 

unnecessary. To produce the correct approach flow to the end blades 

(B1 and B7), the amount of flow escaping around the two ends of the 

cascade was controlled with the flaps shown in Fig. 1. The flap 

positions were set using a wool tuft upstream of each blade to check 

that the incoming flow approached the stagnation points with the 

correct angle. The inlet flow angle at the center of the cascade was 

also checked with a three-hole pressure probe and found to agree with 

the design angle to within about 2° of uncertainty. At high Reynolds 

numbers, the approach velocity to the middle four passages was 

measured to be uniform to within 6%, and the difference between any 

two adjacent passages was within 3%. At low Reynolds numbers, 

slightly more variation was observed, but the approach velocity to the 

middle two passages still agreed to within 5%. Good periodicity at 

high Reynolds numbers was also observed in the exit flow from the 

cascade, as evidenced by suction side velocity profiles acquired near 

the trailing edge of blades B2-B6, and by total pressure loss surveys, 

which are shown below. At low Reynolds numbers, when significant 

separation bubbles were present, the periodicity was not as good due 

to suppression of the separation bubble thickness on the blades closest 

to the tailboard. This deviation from periodicity is considered 

acceptable for the present facility, since its intended purpose is for the 

study of flow control, which if successful should suppress separation 

on all blades, thereby restoring periodicity even at low Reynolds 

numbers. 

tailboard

flap

flap

screen inlet

outlet

periodic

periodic

B1

B2

B3

B4

B5

B6

B7

Fig. 1 Schematic of a linear cascade of 7 airfoils (Volino 

[18]) with boundary conditions and computational domain 

used in present study, shown in dashed lines 

COMPUTATIONAL METHOD 

 The CFD predictions were performed with the numerical software 

tool FLUENT [23]. Predictions with three turbulence models were 

compared with the experimental results of Volino [18]. These models 

are described below. The computational domain consisted of two 

channels and the airfoil in the middle. It is shown on Fig. 1 in dashed 

lines. The boundary conditions applied along the sides of the domain 

are periodic, except along the airfoil surfaces, which are walls. Noslip boundary conditions were applied for all walls. 

 The inlet boundary condition was a prescribed uniform velocity, and 

is described in more detail below. The inlet into the two channels is 

located at 1.9 Cx upstream of the airfoil leading edge in the flow 

direction. At the outlet zero gage pressure was applied. The exit was 

located at 3.8 Cx downstream of the airfoil trailing edge in the flow 

direction. Different exit locations were tested to insure that 3.8 Cx was 

far enough downstream to achieve independent results (as indicated by 

pressure coefficients, pressure losses downstream of the cascade and 

velocity profiles on the airfoil) through the passages. 

 Unsteady calculations were performed for all cases. Convergence 

was established when: 1) residuals reduced to a value of 10-4, 2) 

velocity at the outlet and pressure on the airfoil suction side settled 

(fluctuations around certain mean values were observed, however 

there were no equal cycles), and 3) the mass imbalance was less than 

0.01 %. After convergence was achieved within each time step (t = 

0.0005 s) with the conditions listed above, time averaged results were 

obtained for 2000 time steps. 

Grid independence study 

 A grid independence study for 2D grids was conducted for Re = 

102,000. The V2F turbulence model (described below) was applied. 

Differences in grids are summarized in Table 2. Grids 2 and 3 showed 

maximum difference in Cp on the suction side of an airfoil less than 

2%, therefore grid 2 was used in this study. 

Table 2 Grid characteristics for grid independence study 

Grid 

number 

Size y+

 on the airfoil 

walls 

Number of 

points on 

airfoil ss 

Number of 

points on 

airfoil ps 

1 91516 2.95 

(range 0.1-8) 

290 240 

2 62469 0.20 

(range 0.01-0.6) 

290 240 

3 312393 0.20 

(range 0.01-0.6) 

1369 933 

Boundary conditions influence study 

 After grid independence was established several inlet velocity 

boundary conditions were tested. First, a uniform inlet velocity was 

applied in the direction of the design inlet flow angle (35°). This 

condition resulted in a slightly higher pressure on the leading section 

of the suction side of the airfoil compared to the experiment, 

indicating that the actual inlet angle could be different from the design 

angle. To investigate this possibility, a 2D inviscid calculation was 

done for the full cascade shown in Fig. 1, including the tailboard and 

flaps. The inlet velocity magnitude and direction were taken from the 

inviscid calculation along a line parallel to the blade row and 1.9 Cx

upstream of the blades in the flow direction, and used to set the inlet 

boundary conditions to the 2 channel domain described above. The 

inlet boundary conditions tested are summarized in Table 3, all were 

tested at Re=102,000. Four different inlet conditions are presented in 

Table 3. The first condition assumed a uniform inlet velocity at the 

design flow angle of 35°. The second condition used the velocity 

profile entering the two channel domain as obtained from the invisicid 

calculation. This condition shows a slight variation in the velocity 

profile at the inlet both in x and  directions with spatial averaged 

velocities Vx= 3.78 and V = 2.32 m/s. Accordingly, the spatial 

averaged inlet flow angle is 31.5°. The third condition used a uniform 

inlet velocity and flow angle (Vx= 3.78 and V = 2.32 m/s and inlet 

flow angle = 31.5°) based on the average values across the inlet of the 

two channel domain from the inviscid calculation. Thus the difference 

between conditions (2) and (3) are in the spatial variation in the inlet 

velocity while the averaged values are the same. The fourth condition 

NASA/CR—2012-217415 43
4 Copyright © 2008 by ASME 

used a uniform inlet velocity and angle based on the spatial averaged 

values across the full cascade (instead of averaging over two channels 

only as in condition (3)) from the inviscid calculation. Therefore the 

inlet velocity variations chosen in this exercise involve different inlet 

flow velocity magnitude and angle based on the design or the inviscid 

CFD results for the 7 blades (see Fig. 1). 

 Boundary condition 4 produced results in better agreement with the 

experimental data and was chosen for the rest of this investigation. 

The deviation of the inlet angle from the design angle is about 2°, 

which is within the uncertainty of the experimental measurement and 

therefore plausible. 

Table 3 Inlet boundary conditions for 2 channels tested, 

Re=102,000 

# Description Vx, 

m/s 

V, 

m/s 

Vmag, 

m/s 

1, 

deg 

1 Design condition 3.65 2.56 4.46 35 

2 Velocity profile from 

cascade simulation taken 

at inlet into 2 channels 

3.78 

(avg) 

2.32 

(avg) 

4.43 

(avg) 

31.5 

(avg) 

3 Velocity from cascade 

simulation averaged 

across inlet into 2 

channels 

3.78 2.32 4.43 31.5 

4 Velocity from cascade 

simulation averaged 

across full cascade at 

streamwise location of 

inlet into 2 channels. 

3.71 2.36 4.4 32.5 

Turbulence Modeling 

 In this study, the k  - sst model of Menter [24], the 

 fv 2 model of Durbin [25], and new Transition-sst (4 eq.) model 

of Menter [21], were compared for separated flow predictions on the 

highly loaded LPT airfoil. The unsteady Reynolds-averaged NavierStokes (URANS) equations were used as the transport equations for 

the mean flow. The governing equations and the detailed description 

of each model have been published elsewhere and therefore are not 

shown in this paper. 

 Shear-stress transport k  model (SKW-sst): This 

model, developed by Menter [24] is similar to the standard k  of 

Wilcox [26], but has an ability to account for the transport of the 

principal shear stress in adverse pressure gradient boundary layers. 

The model is based on Bradshaw's [27] assumption that the principal 

shear stress is proportional to the turbulent kinetic energy, which is 

introduced into the definition of the eddy-viscosity. These features 

make the SST k  model more accurate and reliable for a wider 

class of flows (e.g., adverse pressure gradient flows, airfoils, transonic 

shock waves) than the standard k  model [23]. 

 fv 2 model (V2F): According to Launder [28], the normal 

stress 

2

v , perpendicular to the local streamline plays the most 

important role to the eddy viscosity. Motivated by this idea, Durbin 

[25] devised a “four-equation” model, known as the 2   vk

model, or  fv 2 model. It eliminates the need to patch models in 

order to predict wall phenomena like heat transfer or flow separation. 

It makes use of the standard k   model, but extends it by 

incorporating the anisotropy of near-wall turbulence and non-local 

pressure strain effects, while retaining a linear eddy viscosity 

assumption. 

 Transition-sst (4 eq.) model: A new correlation-based 

transition model was proposed by Menter et al. [21]. This model is 

based on two transport equations. The intermittency transport 

equation is used to trigger the transition onset. The transport equation 

for the transition momentum thickness Reynolds number (Ret) is used 

to capture non-local effects of freestream turbulence intensity and 

pressure gradient at the boundary layer edge. Outside the boundary 

layer the transport variable was forced to follow the value of Ret

given by correlations. Those two equations were coupled with the 

shear stress transport turbulence model (SST). This model has recently 

become available in Fluent code.

RESULTS 

Re =25,000 

 Figure 2 shows the pressure coefficient plotted versus s/Ls along the 

suction and pressure surfaces for the turbulence models tested and Re 

= 25,000. Notice the abscissa is the distance along the blade surface 

normalized with respect to the suction surface length. The 

experimental data shows that the Cp values have a constant value on 

the downstream part of the suction side. This plateau indicates that the 

boundary layer has separated and that the boundary layer never 

reattaches (i.e. the separation bubble bursts); this of course refers to 

the time-averaged bubble as the bubble will probably burst 

sporadically. All turbulence models predict the pressure coefficient 

reasonably well with some deviation near the leading edge. The 

deviation could be partially attributed to uncertainty in the 

measurements and the inlet velocity profile as discussed above. All 

the models do well in predicting the separation location and the failure 

of the flow to reattach, as seen in the experiment. 

 Figure 3 shows the total pressure loss coefficient, , plotted versus 

dimensionless distance (/L) at a location 0.63 Cx downstream of the 

cascade. The definition of used in this study is the same as the one 

used in earlier work by Volino [18] and Bons et al. [29]. The 

coordinate  indicates the distance in the direction perpendicular to the 

axial chord. The normalizing quantity L is the blade spacing. The 

origin, =0, corresponds to the location directly downstream of the 

trailing edge of the center blade in the direction of the exit design flow 

angle. At low Reynolds numbers the burst separation bubble results in 

high losses and forces the peaks about 0.35L toward the pressure side 

of each passage. The peaks become noticeably smaller moving from 

B5 to B3, indicating the effect of the tailboard in reducing the 

separation bubble thickness. Due to the lack of periodicity in the 

experiment, the predicted loss coefficient is not expected to agree 

precisely with the experiment. The prediction is based on periodic 

boundary conditions and is not influenced by tailboard effects, so it 

should show higher losses and possibly lower flow turning (peaks 

shifted more to the left in the figure). This is indeed the case, as 

shown in Fig. 3. 

 Velocity profiles for the six suction surface measurement stations on 

blade B4 documented in the experiment are shown in Fig. 4 for the 

nominal Re=25,000 case. The figure shows the distance from the wall 

normalized on the suction surface length plotted against the local mean 

velocity normalized on the nominal exit velocity, Ue. The boundary 

layer has just separated at the first measurement station and the 

separation bubble grows larger at the downstream stations. 

NASA/CR—2012-217415 44
5 Copyright © 2008 by ASME 

0.00

0.50

1.00

1.50

2.00

2.50

0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0

s/Ls

Cp

Exp.

SKW-sst

Trans-sst

V2F

Fig. 2 Cp profiles, Re=25,000 

Fig. 3 Pressure loss coefficients at 0.63 Cx downstream of 

cascade, Re=25,000 

Fig. 4 Mean velocity profiles, Re = 25,000 

Fig. 5 Comparison of Transition-sst and experimental u′/Ue

profiles, Re= 25,000 

 

The boundary layer does not reattach. The velocity profiles at the six 

stations along the suction surface are predicted reasonably well by all 

models with the Transition model doing better overall. The prediction 

in the near wall region is different than the data since the 

measurements were done using hot-wire anemometry and therefore 

can not register negative velocities when separation occurs. 

 Figure 5 shows u′/Ue profile versus y/Ls at the six different stations 

along the suction side. u′ was obtained from CFD (using the Transition 

model only because of its best overall prediction) from the kinetic 

energy of turbulence and assuming an isotropic flow field (i.e. u′ = v′

=w′). Despite this assumption the CFD shows reasonable agreement 

with the data. The CFD shows a peak in the u′ profile that moves 

away from the wall as one travels from station (1) to (6). This peak 

value will be utilized (as will be shown later) to predict the start of 

transition. It is interesting that the experimental data show a very large 

peak located in the shear layer over the separation bubble which 

reaches a dimensionless magnitude of about 0.2. 

 At this Re (=25,000) the Transition model seems to perform the best 

(compared to SKW and V2F). 

Re =100,000 

 Figure 6 shows the pressure coefficient plotted versus s/Ls along the 

suction and pressure surfaces for the turbulence models tested and Re

= 100,000. All turbulence models predict the pressure coefficient very 

well except the Transition model, which shows under-prediction 

downstream from the suction peak. One possible explanation for this is 

that at s/Ls = 0.6 transition starts (as discussed below). The mixing 

associated with transition will tend to promote reattachment, which 

would result in a drop in the pressure coefficient. The simulation may 

be over predicting this tendency toward reattachment in this case, 

although the velocity profiles shown below do not indicate 

reattachment. 

NASA/CR—2012-217415 45
6 Copyright © 2008 by ASME 

0.00

0.50

1.00

1.50

2.00

2.50

0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0

s/Ls

Cp

Exp.

SKW-sst

Trans-sst

V2F

Fig. 6 Cp profiles, Re=100,000 

Fig. 7 Pressure loss coefficients at 0.63 Cx downstream of 

cascade, Re=100,000 

Fig. 8 Mean velocity profiles, Re = 100,000 

Fig. 9 Comparison of Transition-sst and experimental u′/Ue

profiles, Re= 100,000 

 Figure 7 shows the total pressure loss coefficient plotted versus 

dimensionless distance at 0.63 Cx downstream of the cascade. The loss 

coefficient was predicted reasonably well by all models with the 

Transition model showing the best agreement. As in the Re = 25,000 

case, the experimental results were not periodic, so precise agreement 

is not expected between the prediction and the data. 

 Figure 8 shows the velocity profiles (normalized with respect to the 

exit free stream velocity) versus y/Ls at the 6 stations along the suction 

side. The velocity profiles are predicted reasonably well by all models 

with the Transition model doing better overall. Again the near wall 

region is different than the data since the measurements were done 

using hot-wire anemometry and therefore can not register negative 

velocities when separation occurs. 

 Figure 9 shows u′/Ue profiles versus y/Ls at 6 stations along the 

suction side. As explained above, u′ was obtained from the CFD 

(using the Transition model) using the kinetic energy of turbulence 

and assuming an isotropic flow field. Again the CFD results show a 

peak in the u′ profile that moves away from the wall as one travels 

from station (1) to (6). Similarly, from the experimental data, the 

separation bubble grows in the streamwise direction, the boundary 

layer does not reattach, and there is a very high peak in the fluctuating 

velocity in the shear layer over the separation bubble. 

 At this Re (=100,000) the Transition model seems to perform the 

best (compared to SKW and V2F). Its one discrepancy is underprediction in the pressure coefficient downstream of the suction peak. 

Re =300,000 

 The experimental data for this case indicate that the boundary layer 

is attached over most of the airfoil. Figure 10 shows the pressure 

coefficient plotted versus s/Ls along the suction and pressure surfaces 

for the turbulence models tested and Re = 300,000. All turbulence 

models predict the pressure coefficient reasonably well including the 

area near the leading edge. The Transition model shows a bump at 

s/Ls = 0.6 indicating a small bubble that appears and then closes 

quickly. Although not as clear in the data, the predicted bubble may 

be correct. The predicted location is between two pressure 

measurement locations in the experiment, so the bubble presence 

would not necessarily be visible in the data. Also, a small bubble at 

this location was clearly noticed in the experimental data at a lower Re

(=200,000). 

NASA/CR—2012-217415 46
7 Copyright © 2008 by ASME 

0.00

0.50

1.00

1.50

2.00

2.50

0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0

s/Ls

Cp

Exp.

SKW-sst

Trans-sst

V2F 

Fig. 10 Cp profiles, Re=300,000 

Fig. 11 Pressure loss coefficients at 0.63 Cx downstream of 

cascade, Re=300,000 

Fig. 12 Mean velocity profiles, Re =300,000 

Fig. 13 Comparison of Transition-sst and experimental u′/Ue

profiles, Re= 300,000 

 Figure 11 shows the total pressure loss coefficient plotted versus 

dimensionless distance at 0.63 Cx downstream of the cascade. The 

loss coefficient was predicted reasonably well by all models. The 

experimental data showed periodic results at this higher Re. Therefore 

the periodic boundary condition applied in the CFD is consistent with 

the experiment. Nevertheless, the magnitude of the pressure loss 

coefficient is over predicted by about 15%. The location of the peaks 

is also shifted to the left of the experimental peaks in Fig. 11. The 

amount of the shift corresponds to about a 4° difference in flow angle. 

 Figure 12 shows the velocity profile (normalized with respect to the 

exit free stream velocity) versus y/Ls at 6 stations along the suction 

side. The velocity profiles are predicted very well by all models. 

Figure 13 shows u′/Ue profiles versus y/Ls at the 6 stations along the 

suction side. The CFD shows good agreement with the data. 

 At this Re (300,000) all models perform very similar with each 

other. The Transition model shows a small bump in the pressure 

coefficient downstream from the suction peak indicating the presence 

of a small bubble at that location. 

Prediction of Transition 

 From the above investigation one can conclude that the Transition 

model shows overall better agreement with the experimental data. 

Therefore in this section this model will be utilized to show how its 

predictions compare with data for the locations of separation and the 

start of transition. 

 Figure 14 shows contours of u′/Ue over the suction side of the airfoil 

overlapped with velocity vectors at 14a) Re=25,000, 14b) Re =100,000 

and 14c) Re = 300,000. On each plot the location of the 1) suction 

peak, 2) six stations used earlier in the velocity comparison with the 

data, 3) separation point and 4) transition start are shown. The CFD 

results of Fig. 14a show that the flow separates (at station (1)) with no 

reattachment, as observed experimentally. The location of transition 

was taken as the location where u′/Ue peaks in the shear layer over the 

separation bubble (between station (3) and (4) and close to (4)). In 

Fig. 14b the CFD data show that the flow separates (at station (1)) 

with no reattachment as observed experimentally. Again the location 

of transition was obtained where u′/Ue peaks in the shear layer 

(between station (1) and (2)). Notice that the suction peak is at about 

the same location as in the case with Re = 25,000. In Fig. 14c the CFD 

data show that with the possible exception of a small bubble appearing 

NASA/CR—2012-217415 47
8 Copyright © 2008 by ASME 

a) 

b) 

c) 

Fig. 14 Contours of u′/Ue, and velocity vectors (for 

Transition model) showing the location of: 1) suction peak, 

2) separation and 3) transition for a) Re = 25,000, b) Re = 

100,000, and c) Re = 300,000 

Fig. 15 Comparison between CFD data (Transition Model) 

and correlation for the start of transition 

at s/Ls = 0.6, the boundary layer is attached, which is consistent with 

the experimental data. Again the location of transition was obtained 

where the u′/Ue peaks (between station (2) and (3)). 

 To quantify the results in Fig. 14, Table 4 shows the numerical 

values of the predicted momentum thickness Reynolds number at the 

suction peak (Rep), the Reynolds number based on the freestream 

velocity at the suction peak and the streamwise distance from the 

suction peak to transition (Rept), and the streamwise locations of the 

suction peak, transition and separation. Table 5 shows corresponding 

measured quantities from the experiment. Note that Rep and ss in 

Table 5 were approximated using a laminar boundary layer calculation 

as explained in Volino [18]. The ranges given for Rept and st/Ls result 

from the finite spacing between measurement stations. The transition 

location is shown again in Fig. 15 along with a correlation from 

Volino and Bohl [30]: 

Rept=8.80 [6.37 – log10(TI2)] Rep

4/3

The agreement between the CFD and experiment shown in Tables 4 

and 5 and Fig. 15 is reasonably good. 

Table 4 CFD results for separation and transition locations 

(Transition Model) 

Re 25,000 100,000 300,000 

Rep 45 87 165 

Rept 11274 20816 73145 

sp/Ls 0.43 0.42 0.47 

st/Ls 0.74 0.59 0.64 

ss/Ls 0.53 0.53 0.6 

CONCLUSIONS 

 The flow over the very high lift L1A airfoil was studied 

experimentally using a linear cascade and reported by Volino [18]. 

Reynolds numbers based on suction surface length and nominal exit 

velocity ranging from 25,000 to 330,000 were considered. The 

experimental data showed that in all cases the laminar suction surface 

boundary layer separated, but at Reynolds numbers greater than 

150,000 the separation bubble was very thin and short, and the 

boundary layer was attached over most of the surface. At lower 

Transition 

Suction Peak Station 

1 

2 

3 

4 

5 

6 

Separation 

Suction Peak 

Transition 

Suction Peak Station 

1 

2 

3 

4 

5 

6 Separation 

Suction Peak 

Transition 

Station 

1 

2 

3 

4 

5 

6 

NASA/CR—2012-217415 48
9 Copyright © 2008 by ASME 

Reynolds numbers the boundary layer separated and never reattached. 

Transition to turbulence occurred in all cases in the shear layer after 

separation. Transition caused immediate reattachment in the high 

Reynolds number cases, but the turbulent shear layer remained 

separated in the low Re cases. 

 Three different CFD URANS models were utilized in this study, the 

SKW-sst, V2F and Transition models. At Re=25,000, the Transition 

model seems to perform the best. At Re=100,000 the Transition model 

again seems to perform the best, although it under-predicts in the 

pressure coefficient downstream of the suction peak. At Re= 300,000 

all models perform very similar with each other. The Transition 

model shows a small bump in the pressure coefficient downstream 

from the suction peak indicating the presence of a small bubble at that 

location. 

 Upon comparing the pressure loss coefficient at 0.63 Cx downstream 

of the cascade, the CFD shows a shift toward the pressure side of the 

passage compared to the data. Further investigation of the cause of 

this shift is needed. 

 Reasonably good agreement was obtained upon comparing the start 

of transition as obtained from CFD (using the Transition model), a 

published correlation and the experimental data. 

Table 5 Experimental Results for separation and transition 

locations 

Re 25,000 100,000 300,000 

Rep 48 96 193 

Rept 12140 

+/-3300 

28340 

+/-6500 

71170 

+/-22600 

sp/Ls 0.44 0.44 0.49 

st/Ls 0.78 

+/-0.094 

0.64 

+/-0.047 

0.64 

+/-0.047 

ss/Ls 0.50 0.50 0.54 

ACKNOWLEDGMENTS 

 This work was sponsored by the National Aeronautics and Space 

Administration. The grant monitor is Dr. Anthony Strasizar of the 

NASA Glenn Research Center. 

REFERENCES 

[1] Mayle, R.E., 1991, “The Role of Laminar-Turbulent Transition in 

Gas Turbine Engines,” ASME Journal of Turbomachinery, 113, 

pp. 509-537 

[2] Howell, R. J., (1999) “Wake - Separation Bubble Interactions in 

Low Reynolds Number Turbomachinery,” Ph.D. dissertation, 

Cambridge University, Cambridge, UK. 

[3] Singh, N., (2005) “A Study of Separated Flow through a LowPressure Turbine Cascade,” MS thesis, University of Cincinnati, 

Cincinnati, OH. 

[4] Volino, R.J., 2002, “Separated Flow Transition Under Simulated 

Low-Pressure Turbine Airfoil Conditions: Part 1 –Mean Flow 

and Turbulence Statistics,” ASME Journal of Turbomachinery, 

124, pp. 645-655. 

[5] Volino, R.J., 2002, “Separated Flow Transition Under Simulated 

Low-Pressure Turbine Airfoil Conditions: Part 2 – Turbulence 

Spectra,” ASME Journal of Turbomachinery, 124, pp. 656-664. 

[6] Mahallati, A., Sjolander, S.A., McAuliffe, B.R., and Praisner, 

T.J., 2007, “Aerodynamics of a Low-Pressure Turbine Airfoil at 

Low-Reynolds Numbers Part 1 -- Steady Flow Measurements, 

ASME Paper GT2007-27347. 

[7] Mahallati, A., and Sjolander, S.A., 2007, “Aerodynamics of a 

Low-Pressure Turbine Airfoil at Low Reynolds Numbers Part 2 --

Blade-Wake Interaction,” ASME Paper GT2007-27348. 

[8] Zoric, T., Popovic, I., Sjolander, S.A., Praisner, T., and Grover, 

E., 2007, “Comparative Investigation of Three Highly Loaded LP 

Turbine Airfoils: Part I -- Measured Profile and Secondary Losses 

at Design Incidence,” ASME Paper GT2007-27537. 

[9] Zhang, X.F., and Howard Hodson, H., 2007, “Effects of Reynolds 

umber and Freestream Turbulence Intensity on the Unsteady 

boundary Layer Development on an Ultra-High-Lift LPT airfoil,” 

ASME Paper GT2007-27274 

[10] Zhang, X.F., Vera, M., Hodson, H., and Harvey, N., 2007, 

“Separation and Transition Control on an Aft-Loaded UltraHigh-Lift LP Turbine Blade at Low Reynolds Numbers: LowSpeed Investigation,” ASME Journal of Turbomachinery, 128, 

pp. 517-527. 

[11] Bohl, D.G. and Volino, R.J., 2006, “Experiments with ThreeDimensional Passive Flow Control Devices on Low-Pressure 

Turbine Airfoils,” ASME Journal of Turbomachinery, 128, pp. 

251-260. 

[12] Volino, R.J., 2003, “Passive Flow Control on Low-Pressure 

Turbine Airfoils,” ASME Journal of Turbomachinery, 125, pp. 

754-764. 

[13] Bons, J.P., Sondergaard, R., and Rivir, R.B., 2001, “Turbine 

Separation Control Using Pulsed Vortex Generator Jets,” ASME 

Journal of Turbomachinery, 123, pp. 198-206. 

[14] Volino, R.J., 2003, “Separation Control on Low-Pressure Turbine 

Airfoils Using Synthetic Vortex Generator Jets,” ASME Journal 

of Turbomachinery, 125, pp. 765-777. 

[15] Huang, J., Corke, T., and Thomas, F., 2003, “Plasma Actuators 

for Separation Control on Low Pressure Turbine Blades,” AIAA 

Paper 2003-1027. 

[16] Praisner, T.J., and Clark, J.P., 2007, “Predicting Transition in 

Turbomachinery – Part 1: A Review and New Model 

Development,” ASME Journal of Turbomachinery, 129, pp. 1-

13. 

[17] Bons, J.P., Hansen, L.C., Clark, J.P., Koch, P.J., and 

Sondergaard, R., 2005, “Designing Low-Pressure Turbine Blades 

With Integrated Flow Control,” ASME Paper GT2005-68962. 

[18] Volino, R.J. "Separated Flow Measurements on a Highly Loaded 

Low-Pressure Turbine Airfoil", ASME Paper GT2008-51445. 

[19] McAuliffe, B.R., and Yaras, M.I., 2007, “Transition Mechanisms 

in Separation Bubbles Under Low and Elevated Freestream 

Turbulence,” ASME Paper GT2007-27605. 

[20] Suzen, Y. B., Huang, P. G., Volino, R.J., Corke, T.C., Thomas, F. 

O., Huang, J, Lake, J. P., King, P.I., 2003, “A Comprehensive 

CFD Study of Transitional Flows in Low-Pressure Turbines 

Under a Wide Range of Operating Conditions,” AIAA Paper 

AIAA2003-3591. 

[21] Menter, F.R., Langtry, R.B., Likki, S.R., Suzen, Y.B., Huang, 

P.G., and Völker, S., 2006, “A Correlation based Transition 

Model using Local Variables Part 1- Model Formulation”, ASME 

Journal of Turbomachinery, 128, pp. 413-422. 

[22] Langtry, R.B., Menter, F.R., Likki, S.R., Suzen, Y.B., Huang, 

P.G., and Völker, S., 2006, “A Correlation based Transition 

Model using Local Variables Part II: Test Cases and Industrial 

Applications,” ASME Journal of Turbomachinery, 128, pp. 423-

444. 

[23] Fluent 6, 2001, Users manual, Fluent Inc., Lebanon NH, USA. 

[24] Menter, F. R., 1994, “Two-Equation Eddy-Viscosity Turbulence 

Models for Engineering Applications,” AIAA Journal, 32(8), pp. 

1598-1605. 

NASA/CR—2012-217415 49
10 Copyright © 2008 by ASME 

[25] Durbin, P.A., 1995, "Separated Flow Computations with the k--

v2 Model," AIAA Journal, 33(4), pp. 659-664. 

[26] Wilcox, D.C., 1998, Turbulent Modeling for CFD, DCW 

Industries, Inc., La Canada, California. 

[27] Bradshaw, P., Ferriss, D.H., and Atwell, N.P., 1967, “Calculation 

of Boundary-Layer Development Using the Turbulent Energy 

Equation,” Journal of Fluid Mechanics, 28(3), pp. 593-616. 

[28] Launder, B.E., and Spalding, D.B., 1974, “The Numerical 

Computation of Turbulent Flows,” Computer Methods in Applied 

Mechanics and Engineering, 3, pp. 269-289. 

[29] Bons, J.P., Pluim, J., Gompertz, K., Bloxham, M., and Clark, J.P., 

2008, “The Application of Flow Control to an Aft-Loaded Low 

Pressure Turbine Cascade with Unsteady Wakes,” ASME Paper 

GT2008-50864. 

[30] Volino, R.J., and Bohl, D.G., 2004, “Separated Flow Transition 

Mechanisms and Prediction With High and Low Freestream 

Turbulence Under Low Pressure Turbine Conditions,” ASME 

Paper GT2004-63360. 

 

NASA/CR—2012-217415 50
1 Copyright © 2008 by ASME 

Proceedings of IMECE2008 

2008 ASME International Mechanical Engineering Congress and Exhibition 

November 2-6, 2008, Boston, Massachusetts, USA

IMECE2008-68776 

EXPERIMENTAL AND COMPUTATIONAL INVESTIGATIONS OF SEPARATION AND 

TRANSITION ON A HIGHLY LOADED LOW-PRESSURE TURBINE AIRFOIL: PART 2 – 

HIGH FREESTREAM TURBULENCE INTENSITY

Ralph J. Volino

Mechanical Engineering Department 

United States Naval Academy 

Annapolis, Maryland 21402-5042 

volino@usna.edu 

Olga Kartuzova and Mounir B. Ibrahim 

Mechanical Engineering Department 

Cleveland State University 

Cleveland, Ohio 44115-2425 

m.ibrahim@csuohio.edu 

ABSTRACT 

 Boundary layer separation, transition and reattachment have been 

studied on a very high lift, low-pressure turbine airfoil. Experiments 

were done under high (4%) freestream turbulence conditions on a 

linear cascade in a low speed wind tunnel. Pressure surveys on the 

airfoil surface and downstream total pressure loss surveys were 

documented. Velocity profiles were acquired in the suction side 

boundary layer at several streamwise locations using hot-wire 

anemometry. Cases were considered at Reynolds numbers (based on 

the suction surface length and the nominal exit velocity from the 

cascade) ranging from 25,000 to 300,000. At the lowest Reynolds 

number the boundary layer separated and did not reattach, in spite of 

transition in the separated shear layer. At higher Reynolds numbers 

the boundary layer did reattach, and the separation bubble became 

smaller as Re increased. High freestream turbulence increased the 

thickness of the separated shear layer, resulting in a thinner separation 

bubble. This effect resulted in reattachment at intermediate Reynolds 

numbers, which was not observed at the same Re under low freestream 

turbulence conditions. Numerical simulations were performed using 

an unsteady Reynolds averaged Navier-Stokes (URANS) code with 

both a shear stress transport k-ω model and a 4 equation shear stress 

transport Transition model. Both models correctly predicted 

separation and reattachment (if it occurred) at all Reynolds numbers. 

The Transition model generally provided better quantitative results, 

correctly predicting velocities, pressure, and separation and transition 

locations. The model also correctly predicted the difference between 

high and low freestream turbulence cases. 

NOMENCLATURE 

Cf skin friction coefficient

Cp ( ) 2

T − ρU/PP2 e , pressure coefficient 

Cx axial chord length 

f frequency

H δ

∗ /θ, shape factor 

Ls suction surface length 

Lφ blade spacing (pitch) 

P pressure 

PS upstream static pressure 

PT upstream stagnation pressure 

PTe downstream stagnation pressure 

PSD power spectral density

Re LU se /ν , exit Reynolds number 

Reθ momentum thickness Reynolds number, U∞θ/ν

s streamwise distance from leading edge along surface 

TI freestream turbulence intensity 

U mean streamwise velocity 

Ue nominal exit freestream velocity, based on inviscid solution 

U∞ local freestream velocity 

u′ time averaged rms streamwise fluctuating velocity 

x axial distance from leading edge 

y distance from wall 

y+ distance from wall in viscous lengths 

Zw Zweifel coefficient 

α1 inlet flow angle 

α2 exit flow angle 

δ

∗ displacement thickness 

φ coordinate along blade spacing, normal to axial chord 

γ intermittency, fraction of time flow is turbulent 

ν kinematic viscosity 

ρ density 

θ momentum thickness 

ψ (PT-PTe)/(PT-PS), total pressure loss coefficient 

ψint total pressure loss integrated over blade spacing 

Subscripts 

p pressure minimum (suction peak) location 

s separation location 

t transition start 

INTRODUCTION 

 Boundary layer separation can lead to partial loss of lift and higher 

aerodynamic losses on low-pressure turbine (LPT) airfoils (e.g. 

Hourmouziadis [1], Mayle [2], and Sharma et al. [3]). As designers 

impose higher loading to improve efficiency and lower cost, the 

Proceedings of IMECE2008 

2008 ASME International Mechanical Engineering Congress and Exposition 

October 31-November 6, 2008, Boston, Massachusetts, USA

IMECE2008-68776

NASA/CR—2012-217415 51
2 Copyright © 2008 by ASME 

associated strong adverse pressure gradients on the suction side of the 

airfoil can exacerbate separation problems, particularly at low 

Reynolds numbers. Prediction and control of suction side separation, 

without sacrifice of the benefits of higher loading, is therefore, crucial 

for improved engine design. 

 Separation and separated flow transition, which can lead to 

boundary layer reattachment, have received considerable attention. 

Studies have included flows over flat plates and airfoils in cascades. 

Some have considered steady inflow conditions, while others have 

included the effect of unsteady wakes. Volino [4] lists some recent 

examples. In general, the strong acceleration on the leading section of 

the airfoil keeps the suction side boundary layer thin and laminar, even 

in the presence of elevated freestream turbulence. When separation 

does occur, it is usually just downstream of the suction peak. If 

transition then occurs in the shear layer over the separation bubble, it 

is typically rapid and causes the boundary layer to reattach [5, 6]. 

Transition is dependent on Reynolds number, freestream turbulence 

level, and the surface roughness conditions upstream of the separation 

point. 

 Building on the understanding developed in earlier studies, higher 

lift airfoils with pressure gradients more resistant to separation have 

been designed, as described by Praisner and Clark [7]. Forward 

loading, for example, makes airfoils more separation resistant by 

extending the adverse pressure gradient on the aft portion of the 

suction side over a longer distance. This reduces the local pressure 

gradient at all locations, making separation less likely. If separation 

does occur, forward loading provides a longer distance along the 

airfoil surface for reattachment. An example is the L1M airfoil used 

by Bons et al. [8]. The L1M is a mid-chord loaded design, and is 

resistant to separation even at very low Reynolds numbers. The L1M 

has 10% higher loading than the “ultra-high lift” airfoils described by 

Zhang and Hodson [9], and 17% higher loading than the Pack B airfoil 

considered in several studies such as Volino [5, 6] and Bons et al. [10]. 

The design calculations indicate that the Zweifel coefficient increases 

from about 1.15 on the Pack B to about 1.35 on the L1M. If the 

definition for the Zweifel coefficient is taken as 

Zw=2 cos2

α2(Lφ/Cx)(tanα1+tanα2) (1) 

as given by Lakshminarayana [11], Zw=1.08 for the Pack B (in 

agreement with the value given by McAuliffe and Sjolander [12]) and 

Zw=1.23 for the L1M. The L1M has the same inlet and exit flow 

angles as the Pack B. 

 Although higher lift without separation is possible, forward loading 

has some disadvantages. As noted by Zhang et al. [13], the longer 

region of turbulent flow on a forward loaded airfoil can lead to 

increased profile losses. Forward loading also creates longer regions 

of strong pressure gradient on the endwalls, which can produce 

stronger secondary flows and losses. It might be possible to produce 

an aft loaded airfoil that was resistant to separation and had low profile 

and secondary loss characteristics over a range of Reynolds numbers if 

flow control were incorporated to suppress separation, as discussed by 

Bons et al. [8]. 

 Bons et al. [14], Volino [15], and Ibrahim et al. [16] studied the 

flow over the L1A airfoil. The L1A was designed at the Air Force 

Research Laboratory (AFRL) and is available on a limited basis to 

U.S. researchers from Clark [17]. It is an aft loaded blade with the 

same flow angles and loading as the L1M. Dimensions of the L1A as 

used in the present study are given in Table 1. 

 Volino [15] showed that with low (0.6%) freestream turbulence and 

steady inflow conditions, the boundary layer separates on the L1A at 

Re≤120,000 and the time averaged separation bubble does not show 

Table 1: Cascade parameters 

Axial 

Chord, Cx

[mm] 

True 

Chord 

[mm] 

Pitch, 

Lφ

[mm] 

Span 

[mm] 

Suction 

side, Ls

[mm] 

Inlet 

flow 

angle 

Exit 

flow 

angle 

134 146 136 724 203 35° 60° 

reattachment (i.e. the time averaged separation bubble bursts). The 

Reynolds number, Re, is based on Ls and the nominal exit velocity 

from the cascade Ue (computed using the inlet velocity and the design 

inlet and exit flow angles). At Re≥200,000 a very small separation 

bubble may be present, but the boundary layer is largely attached and 

very high lift is achieved. Transition occurred in all cases, even at Re

as low as 25,000. The L1A results contrast with those of earlier 

studies on airfoils such as the Pack B. At intermediate Re on the Pack 

B, transition in the shear layer over the separation bubble induced 

almost immediate reattachment. With increasing Re, or with elevated 

freestream turbulence, transition moved upstream and the separation 

bubble became progressively smaller. Only at the very lowest Re, 

where the shear layer remained laminar, did the separation bubble 

burst. As shown in Volino [15], the adverse pressure gradient is 

stronger on the L1A than on the Pack B, and this prevented 

reattachment, even with the mixing caused by a turbulent shear layer. 

Separation without reattachment on the L1A resulted in a 20% 

reduction in lift and a factor of 7 increase in total pressure loss. Bons 

et al. [14] reported similar results, although some differences were 

present at intermediate Re, possibly due to the higher (3%) freestream 

turbulence intensity in their study. The Bons et al. [14] work also 

included the effect of upstream wakes and flow control. 

 Ibrahim et al. [16] computed the flow over the L1A airfoil, 

matching the conditions of Volino [15]. They used the commercial 

code Fluent [18] with three of its turbulence and transition models. 

These were the shear-stress transport k-ω model of Menter [19] 

(SKW-sst), the v

2

-f model of Durbin [20], and the transition model of 

Menter et al. [21] (Transition-sst) which includes a shear-stress 

transport model. All three models correctly predicted separation 

without reattachment at Re=25,000 and 100,000 and attached flow at 

Re=300,000. At the highest Re all three models did well and predicted 

about the same pressure profile on the airfoil, total pressure loss 

profiles, and velocity profiles at several streamwise locations in the 

suction side boundary layer. At the lower Re, all models did 

reasonably well predicting the pressure profiles, but the Menter et al. 

[21] transition model was generally better at predicting the velocity 

profiles, particularly when the boundary layer was separated. 

 The present study extends the results of Volino [15] and Ibrahim et 

al. [16] to cases with higher inlet freestream turbulence. Elevated 

freestream turbulence is intended to more closely match engine 

conditions. The results of the present study and those of Volino [15] 

will serve as baseline cases for future work in which unsteady wakes 

and separation control will be added. 

EXPERIMENTAL FACILITY AND MEASUREMENTS

 Experiments were conducted in a closed loop wind tunnel with a 

seven blade linear cascade in one corner of the loop, as shown in Fig. 

1. Details are available in Volino [15]. A fine screen located 5.3 Cx

upstream of the center airfoil and positioned perpendicular to the flow 

was used by Volino [15] to produce uniform inlet flow with low 

freestream turbulence. In the present study, this screen was replaced 

by a coarse grid, consisting of a 1.5 mm thick sheet metal plate with 

19 mm square holes spaced 25.4 mm apart, center to center, in both 

directions. In a plane perpendicular to the inlet flow and 1.7 Cx

upstream of the center blade, the grid produced uniform flow with a 

turbulence intensity, TI, of 6.0% in the streamwise component and 

NASA/CR—2012-217415 52
3 Copyright © 2008 by ASME 

tailboard

flap

flap

grid inlet

outlet

periodic

periodic

B1

B2

B3

B4

B5

B6

B7

Fig. 1 Schematic of linear cascade and computational 

domain (dashed line) 

4.2% in the cross stream components, for an overall intensity of 4.9%. 

The streamwise component was also measured at the inlet plane of the 

cascade, near the center blade, where it had decayed to 4%. The 

integral length scale of the freestream turbulence is 0.12 Cx in the 

streamwise direction and 0.04 Cx in the other directions. 

 The center blade, designated B4 in Fig. 1, contains pressure taps 

near the spanwise centerline. As described in Volino [15], a tailboard, 

was needed to produce the correct exit flow angle from the cascade. 

Its position was set to produce periodicity at high Reynolds numbers. 

The flaps shown in Fig. 1 were used to control the amount of flow 

escaping around the two ends of the cascade and produce the correct 

approach flow. Figure 1 also shows the computation domain (dashed 

line) with boundary conditions used in the CFD analysis. 

Measurements 

 Pressure surveys were made using a pressure transducer (0-870 Pa 

range Validyne transducer). Stagnation pressure was measured with a 

pitot tube upstream of the cascade. Static pressure taps were located in 

the center blade as noted above. The uncertainty in the suction side 

pressure coefficients was 0.07. Most of this uncertainty was due to 

bias error. Stochastic error was minimized by averaging pressure 

transducer readings over a 10 second period. 

 A four component traverse with three linear stages and one rotating 

stage was located in the wind tunnel downstream of the cascade. The 

traverse produced an acceptably low blockage when it was located at 

least two axial chord lengths downstream of the cascade. Variations in 

the wind tunnel velocity were less than 2% as the traverse was moved 

to various positions. The traverse was used to hold and move probes 

for velocity and downstream pressure measurements. 

 Total pressure losses were documented using a Kiel probe traversed 

across three blade spacings, 0.63 Cx downstream of the cascade. To 

compute the mass flow averaged pressure loss, a velocity profile was 

also acquired along the same line as the total pressure surveys using a 

single sensor hot-film probe. 

 Velocity profiles on the suction surface were measured at the six 

streamwise stations listed in Table 2. All stations are downstream of 

the inviscid pressure minimum at s/Ls=0.49. Profiles were measured 

near the spanwise centerline of airfoil B4 with a hot-wire anemometer 

(AA Lab Systems model AN-1003) and a single sensor hot-film probe 

(TSI model 1201-20). The sensor diameter is 51 μm, and the active 

length is 1.02 mm. At each measurement location, data were acquired 

for 26 seconds at a 20 kHz sampling rate (219 samples). All raw data 

were saved. The high sampling rate provides an essentially continuous 

signal, and the long sampling time results in low uncertainty in both 

Table 2: Velocity profile measurement stations 

Station 1 2 3 4 5 6 

s/Ls 0.53 0.59 0.69 0.78 0.88 0.97 

x/Cx 0.65 0.72 0.80 0.86 0.92 0.97 

statistical and spectral quantities. Data were acquired at 40 wall 

normal locations in each profile, extending from the wall to the freestream, with most points concentrated in the near wall region. The 

probe was positioned as close to tangent to the airfoil surface as 

possible at each station, such that the probe body extended 

downstream of the sensor and the direction of the traverse was within 

5° of normal to the surface. In most cases the closest point in each 

profile was within about 0.2 mm of the wall, which compares to 

boundary layer thicknesses ranging from 1.3 mm to 37 mm. 

 Flow direction in a separation bubble cannot be determined with a 

single-sensor hot-wire, so measurements within the bubble provide 

only a rough estimate of the mean and fluctuating velocity magnitude. 

With the exception of the separation bubbles, the uncertainty in the 

mean velocity is 3-5% except in the very near wall region, where nearwall corrections (Wills [22]) were applied to the mean velocity. 

Uncertainties in the momentum and displacement thicknesses 

computed from the mean profiles are 10%. Uncertainty in the shape 

factor, H, is 8%. Local skin friction coefficients were computed from 

the near wall mean velocity profiles using the technique of Volino and 

Simon [23]. This technique accounts for streamwise pressure gradient 

effects on the mean profile. The uncertainty in Cf is 8%. The 

uncertainty in the fluctuating streamwise velocity is below 10%. 

 Pressure surveys and velocity profiles were acquired at nominal 

Re=25,000, 50,000, 100,000, 150,000, 200,000, and 300,000. The 

Reynolds number, as defined above, is based on the suction surface 

length and the nominal cascade exit velocity. The corresponding 

Reynolds numbers based on the cascade inlet velocity and the axial 

chord length ranged from 10,000 to 120,000. 

NUMERICAL SIMULATIONS 

 The turbulence models and computational method used in the 

present study are described briefly here and more completely in 

Ibrahim et al. [16]. An unsteady Reynolds averaged Navier-Stokes 

(URANS) code was used. Sufficient time steps were computed so that 

the flow converged to a steady mean. Results were then time averaged 

based on 2000 time steps computed after convergence. Ibrahim et al. 

[16] considered several turbulence models and selected the three most 

promising noted above (SKW-sst, v

2

-f and Transition-sst) for extended 

study. Of these, the SKW-sst and Transition-sst performed best and 

are used in the present study. The Transition-sst model is a 4 equation 

model and includes an intermittency transport equation. It has been 

validated against experimental data for turbomachinery applications 

(Langtry et al. [24]). 

 The computational domain consisted of the two passages between 

blades B3 and B5 in Fig. 1, with blade B4 in the middle. The 

boundary conditions on the sides of the passages were periodic with 

the exception of the pressure side of blade B3 and the suction side of 

blade B5, which were solid walls. The exit boundary was located 3.8 

Cx downstream of the trailing edges in the flow direction. Zero gage 

pressure was specified at the exit. As described in Ibrahim et al. [16], 

exit locations extending farther downstream were tested to insure the 

results were independent of the location chosen. A uniform velocity 

inflow condition was specified 1.9 Cx upstream of the blade leading 

edges in the flow direction. The inlet flow angle was set to 33°

 based 

on an inviscid calculation of the full cascade shown in Fig. 1. This 

angle agreed with the experimentally measured inlet angle to within 

the experimental uncertainty. The experimental value matched the 35°

NASA/CR—2012-217415 53
4 Copyright © 2008 by ASME 

design inlet angle. Ibrahim et al. [16], in the low TI cases, showed that 

calculated Cp results with the 33°

 angle matched the experimental 

results slightly better than calculations with a 35°

 angle specified, but 

the differences were small and within the experimental uncertainty. 

 The computational grid was two-dimensional, consisting of 62469 

nodes with the closest points to the airfoil walls between y

+=0.01 and 

0.6, with an average of y+=0.2. Ibrahim et al. [16] also considered a 

finer grid to establish grid independence. Simulations were done for 

Reynolds numbers of 25,000, 100,000 and 300,000. The freestream 

turbulence intensity was set to 5% at the inlet to the computational 

domain to match the experiments. 

RESULTS 

Pressure Profiles 

 Pressure profiles for all cases are shown in Fig. 2. The inviscid 

profile for the L1A airfoil, computed using a panel code, is shown for 

comparison. For Re≥50,000, the data generally agree with the inviscid 

profile, indicating that the boundary layer is attached over most of the 

airfoil. The data lie somewhat above the inviscid line on the suction 

side at s/Ls between 0.6 and 0.8, indicating a separation bubble 

followed by reattachment. The size of the bubble appears to decrease 

as Re increases. At Re=25,000 the suction peak is suppressed and the 

Cp values have a constant value on the downstream part of the suction 

side, indicating a separated boundary layer that does not reattach. The 

pressure side Cp values show less variation with Reynolds number, but 

in the Re=25,000 case Cp is slightly higher than in the other cases at 

s/Ls=0.25 and near the trailing edge. This suggests more of a leading 

edge separation bubble on the pressure side and higher velocity near 

the trailing edge when the burst suction side separation bubble forces 

fluid toward the pressure side. 

 The present Cp results agree with the low TI cases of Volino [15] 

for Re≥200,000. At Re between 50,000 and 150,000, the high TI

causes reattachment, whereas the flow remained separated with low 

TI. The high and low TI results bracket the results of Bons et al. [14], 

who had an intermediate TI level. 

 The lift on the airfoil can be determined by integrating the 

difference between the suction and pressure side Cp values along the 

axial direction. The result is shown as a function of Reynolds number 

in Fig. 3. Also shown are the low TI cases of Volino [15] for 

comparison. In the high TI cases, the lift drops about 10% between the 

high Re cases and the Re=25,000 case due to the burst separation 

bubble. In agreement with the difference in reattachment, the lift is 

higher for the high TI cases than the low TI at Re below 200,000. 

Total Pressure Losses 

 The loss coefficient, ψ, is shown for all cases in Fig. 4. The 

coordinate φ indicates the distance in the direction perpendicular to the 

axial chord. The normalizing quantity Lφ is the blade spacing. The 

origin, φ=0, corresponds to the location directly downstream of the 

trailing edge of the center blade in the direction of the exit design flow 

angle. At the two highest Reynolds numbers the losses are low, and 

the location of the loss peaks are in the expected positions downstream 

of the airfoils. This indicates that the actual flow angle is essentially 

equal to the design angle. The peaks downstream of blades B3-B5 are 

similar to each other, indicating periodicity in the cascade. These 

results agree with those of Volino [15] for the corresponding low TI

cases. The loss coefficient is low between the peaks, although not as 

close to zero as in the low TI cases. This suggests some loss 

associated with the decay of the freestream turbulence. At 

Re=100,000 and 150,000 the loss peaks increase slightly, particularly 

on the left side of each peak. At Re=50,000 the loss peak has 

increased further, and is about 40% higher than in the highest Re

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

Re=25,000

Re=50,000

Re=100,000

Re=150,000

Re=200,000

Re=300,000

Inviscid

Fig. 2 Cp profiles 

0 0.5 1 1.5 2 2.5 3 3.5

x 10

5

0.9

1

1.1

1.2

1.3

Re

∫(Cpsuction−Cppressure)dx/Cx

Low TI

High TI

Fig. 3 Lift based on integrated Cp profile 

−1.5 −1 −0.5 0 0.5 1 1.5 0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

φ/Lφ

ψ

B3

B4 B5

B6

Re=25,000

Re=50,000

Re=100,000

Re=150,000

Re=200,000

Re=300,000

Fig. 4 Total pressure loss coefficient at 0.63 Cx

downstream of cascade 

cases. The increasing loss at lower Re is consistent with the increasing 

size of the separation bubble at lower Re, shown in Fig. 2. The 

separation bubble would tend to increase the suction side boundary 

layer thickness after reattachment, which would increase losses and 

NASA/CR—2012-217415 54
5 Copyright © 2008 by ASME 

reduce the flow turning, moving the wake to the left in Fig. 4. At 

Re=25,000, the burst separation bubble results in much higher losses, 

and forces the peaks about 0.35Lφ toward the pressure side of each 

passage. This shift corresponds to a 30° change in the exit flow angle. 

The reduction in flow turning is consistent with the lower lift shown in 

Fig. 3. The Re=25,000 results match those of the corresponding low 

TI case in Volino [15]. The peaks in the Re=25,000 case become 

noticeably smaller moving from B6 to B3, which results from the 

tailboard reducing the separation bubble thickness on the closer 

blades. As discussed in Volino [15], the tailboard was set for 

periodicity in the high Re cases. In cases with burst separation 

bubbles, the periodicity is lost. This is considered acceptable for the 

experiment, since the eventual goal is separation control, and if the 

separation can be suppressed, periodicity will be restored. 

 The integrated loss around the center blade is computed as 

= ∫ ∫ − −

2/L

2/L

2/L

2/L

int Ud Ud

φ

φ

φ

φ

ψ φψ φ (2) 

and is shown in Fig. 5 as a function of Reynolds number. Losses 

increase with decreasing Re, rising by about a factor of 2 between the 

highest and lowest Reynolds numbers. This difference is smaller than 

that observed in the low TI cases. 

 The turbulence associated with the airfoil wakes is shown in Fig. 6. 

The rms fluctuating streamwise velocity at 0.63 Cx downstream of the 

trailing edge is normalized on the nominal exit velocity for each case. 

At the higher Reynolds numbers the boundary layer turbulence has 

decayed to about 4% of Ue., matching the low TI results of Volino 

[15]. A double peak is apparent, with the left and right peaks 

corresponding the suction and pressure side boundary layers 

respectively. Between blades the turbulence intensity is at the 

background level of the wind tunnel. As Re drops to 100,000, there is 

a 20% rise in the wake turbulence, corresponding to the rise in the 

pressure loss in Fig. 4. Dropping to Re=50,000 the wake becomes 

much wider, and there is a large rise in the turbulence, particularly on 

the suction side of the wake, where the peak is about double the high 

Re value. There is some loss of periodicity at Re=50,000, with the 

turbulence rising from B3 to B6. The same rise is observed in the 

pressure loss peaks of Fig. 4, suggesting that the separation bubble is 

large enough that the tailboard is affecting it at Re=50,000. The 

turbulence level is still only about 40% of the value in the low TI

Re=50,000 case, where the separation bubble burst. At Re=25,000 the 

turbulence level is much higher, and in agreement with the loss 

coefficients, the peaks are shifted toward the pressure side of the 

passages. The peak values at Re=25,000 were about 50% higher in the 

corresponding low TI case. 

Velocity Profiles 

 Velocity profiles for the six suction surface measurement stations 

are shown in Fig. 7 for the Re=25,000 case. The low TI profiles of 

Volino [15] are shown for comparison. The top row in the figure 

shows the distance from the wall normalized on the suction surface 

length plotted against the local mean velocity normalized on the 

nominal exit velocity, Ue. The boundary layer has just separated at the 

first measurement station and the separation bubble grows larger at the 

downstream stations. The boundary layer does not reattach. The 

second row in Fig. 7 shows the rms streamwise fluctuating velocity, u′, 

normalized with Ue. There is a very large peak located in the shear 

layer over the separation bubble which reaches a dimensionless 

magnitude of about 0.22. The high and low TI cases are similar, with 

0 0.5 1 1.5 2 2.5 3 3.5

x 10

5

0

0.1

0.2

0.3

0.4

0.5

0.6

Re

ψint

Low TI

High TI

Fig. 5 Integrated total pressure loss coefficient for center 

blade as function of Re

−1.5 −1 −0.5 0 0.5 1 1.5 0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

0.2

φ/Lφ

u′/Ue

B3 B4 B5

B3

B4

B5

B6

Re=25,000

Re=50,000

Re=100,000

Re=150,000

Re=200,000

Re=300,000

Fig. 6 Rms fluctuating streamwise velocity at 0.63 Cx

downstream of cascade 

a thick separation bubble in both cases and strong fluctuations in the 

shear layer over the bubble. The boundary layer thickness is nearly 

equal in the two cases. The shear layer, however, is noticeably thicker 

in the high TI case, extending closer to the wall. The thicker shear 

layer suggests stronger mixing closer to the wall, which would tend to 

promote reattachment. Although reattachment did not occur at 

Re=25,000, a thicker shear layer may help explain differences between 

the high and low TI cases at slightly higher Re. The third row in Fig. 7 

shows the intermittency, γ. The intermittency is the fraction of time 

the flow is turbulent. It was determined at each measurement location 

based on the instantaneous streamwise velocity signal, using the 

technique described in Volino et al. [25]. Turbulent flow is defined 

here to include a range of large and small scale eddies. A boundary 

layer or shear layer may have significant u′ fluctuations but still be 

considered non-turbulent if these fluctuations are induced by an 

external source such as freestream turbulence or are associated with 

instability in a narrow frequency band. Transition to turbulence is 

characterized by the appearance of broadband fluctuations. In the 

intermittency processing routine, the velocity signal is high pass 

filtered, and the appearance of high frequency fluctuations is used to 

distinguish between turbulent and non-turbulent flow. The results in 

Fig. 7 indicate that the shear layer begins transition between stations 3 

and 4. 

 Figure 8 presents another view of the transition process using 

turbulence spectra. The spectra are computed from the fluctuating 

NASA/CR—2012-217415 55
6 Copyright © 2008 by ASME 

0 0 0 0 0 0 1 0

0.1

0.2

U/Ue

y/Ls

a) 1 2 3 4 5 6

0 0 0 0 0 0 0.1 0

0.1

0.2

u′/Ue

y/Ls

b) 1 2 3 4 5 6

0 0 0 0 0 0 1 0

0.1

0.2

γ

y/Ls

c) 1 2 3 4 5 6

Low TI

High TI

Fig. 7 Profiles for Re=25,000 case: 

(a) mean velocity, (b) u′/Ue, (c) intermittency 

10−1 100 101 102

10−10

10−8

10−6

10−4

10−2

100

f⋅Ls/Ue

PSD/(Ls

⋅Ue

)

s/Ls=0.531

0.594

0.688

0.781

0.875

0.969

Fig. 8 Turbulence spectra for Re=25,000 case 

velocity signal acquired at the location of peak u′ at each measurement 

station shown in Fig. 7. Frequencies are resolved from 4.88 Hz to 10 

kHz in 4.88 Hz increments using 4096 point Fast Fourier Transforms 

to compute the spectra. The frequencies and power spectral density 

(PSD) are normalized using Ls and Ue. There is a clear rise in the 

spectra between the second and third measurement stations, reaching a 

final high value by station 4. The spectra indicate that transition is in 

its early stages at station 3 and well underway or possibly complete by 

station 4. This agrees with the rise in intermittency between stations 3 

and 4 in Fig. 7, although the spectra suggest the intermittency should 

reach a fully turbulent value of 1. The narrow range of scales resulting 

from the low velocity of the Re=25,000 case is believed to result in 

artificially low γ values from the intermittency processing routine. At 

higher Re, the range of scales is broader, and γ is believed to be 

correctly computed. 

 The spectra of Fig. 8 show a peak at stations 2 and 3 at a 

dimensionless frequency of 10. This peak suggests a shear layer 

instability which likely initiates transition. Although less distinct, 

spectral peaks are visible just before or just after transition at higher 

Reynolds numbers in the figures below. Volino [6, 15] observed 

similar peaks in the low TI cases and in the boundary layer on the Pack 

B airfoil and associated them with Tollmien-Schlicthing (TS) waves. 

The TS waves were believed to form in the attached boundary layer 

0 0 0 0 0 0 1 0

0.1

0.2

U/Ue

y/Ls

a) 1 2 3 4 5 6

0 0 0 0 0 0 0.1 0

0.1

0.2

u′/Ue

y/Ls

b) 1 2 3 4 5 6

0 0 0 0 0 0 1 0

0.1

0.2

γ

y/Ls

c) 1 2 3 4 5 6

Low TI

High TI

Fig. 9 Profiles for Re=50,000 case: 

(a) mean velocity, (b) u′/Ue, (c) intermittency 

10−1 100 101 102

10−10

10−8

10−6

10−4

10−2

100

f⋅Ls/Ue

PSD/(Ls

⋅Ue

)

s/Ls=0.531

0.594

0.688

0.781

0.875

0.969

Fig. 10 Turbulence spectra for Re=50,000 case 

and grow in the separated shear layer until they became large enough 

to induce transition. The frequencies of the peaks in the present high 

TI cases can similarly be associated with TS frequencies. KelvinHelmholtz rollup of the shear layer, as observed by Stieger and 

Hodson [26], and Roberts and Yaras [27], is another possible 

explanation for the peaks. Hughes and Walker [28], Volino [6], and 

others have noted that in adverse pressure gradient cases, transition 

does not occur through a pure bypass mode, even in the presence of 

elevated freestream turbulence. 

 The velocity profiles for the Re=50,000 case are shown in Fig. 9. 

The boundary layer appears to have just separated at station 2, is 

separated at station 3, and is reattaching by station 4. There is a clear 

difference from the low TI case, which shows a burst separation 

bubble. Transition is underway in the shear layer at station 3 and 

complete by station 4, which likely promotes the reattachment. The 

spectra in Fig. 10 agree with the intermittency, showing a rise between 

stations 2 and 3, and fully turbulent values by station 4. The delay in 

separation relative to the low TI case results in a thinner separation 

bubble, which makes reattachment possible when transition begins. 

 Figure 11 shows the velocity profiles for the Re=100,000 case. 

Separation appears to be delayed relative to the Re=50,000 case to 

about station 3. Transition moves upstream and is complete by station 

3. The result is a near elimination of the separation bubble and a 

NASA/CR—2012-217415 56
7 Copyright © 2008 by ASME 

0 0 0 0 0 0 1 0

0.1

0.2

U/Ue

y/Ls

a) 1 2 3 4 5 6

0 0 0 0 0 0 0.1 0

0.1

0.2

u′/Ue

y/Ls

b) 1 2 3 4 5 6

0 0 0 0 0 0 1 0

0.1

0.2

γ

y/Ls

c) 1 2 3 4 5 6

Low TI

High TI

Fig. 11 Profiles for Re=100,000 case: 

(a) mean velocity, (b) u′/Ue, (c) intermittency 

10−1 100 101 102

10−10

10−8

10−6

10−4

10−2

100

f⋅Ls/Ue

PSD/(Ls

⋅Ue

)

s/Ls=0.531

0.594

0.688

0.781

0.875

0.969

Fig. 12 Turbulence spectra for Re=100,000 case 

significantly thinner boundary layer downstream. The maximum u′/Ue

at the last measurement station is 65% of its level in the Re=50,000 

case. The thinner boundary layer and lower turbulence correspond to 

the drop in total pressure loss and wake turbulence between the 

Re=50,000 and 100,000 cases in Figs. 4-6. Figure 12 shows the 

spectra for the Re=100,000 case. The rise between stations 2 and 3 

indicates transition in agreement with the γ profiles of Fig. 11.

 Velocity profiles and spectra for the Re=200,000 case are shown in 

Figs. 13 and 14. The separation bubble has been effectively 

eliminated. This reduces the boundary layer thickness and fluctuating 

velocity levels relative to the Re=100,000 case. The boundary layer 

also remained attached in the low TI Re=200,000 case, and the profiles 

in Fig. 13 are in close agreement for the high and low TI cases. 

Profiles and spectra for the Re=300,000 case are not shown, but with 

the exception of a slightly thinner boundary layer, appear the same as 

those for the Re=200,000 case. 

 The shape factor, H, and skin friction coefficient Cf are good 

indicators of the state of the boundary layer with respect to separation 

and transition. The shape factor is shown Fig. 15. At the first 

measurement stations, H=1.8 at Re=300,000, indicating an attached 

boundary layer. At Re=25,000, H=3.6, indicating the boundary layer 

is separated or on the verge of separation. With low TI, H at station 1 

was above 2.5 at all Re. Particularly in the high Re cases, high TI

0 0 0 0 0 0 1 0

0.02

0.04

0.06

0.08

U/Ue

y/Ls

a) 1 2 3 4 5 6

0 0 0 0 0 0 0.1 0

0.02

0.04

0.06

0.08

u′/Ue

y/Ls

b) 1 2 3 4 5 6

0 0 0 0 0 0 1 0

0.02

0.04

0.06

0.08

γ

y/Ls

c) 1 2 3 4 5 6

Low TI

High TI

Fig. 13 Profiles for Re=200,000 case: 

(a) mean velocity, (b) u′/Ue, (c) intermittency 

10−1 100 101 102

10−10

10−8

10−6

10−4

10−2

100

f⋅Ls/Ue

PSD/(Ls

⋅Ue

)

s/Ls=0.531

0.594

0.688

0.781

0.875

0.969

Fig. 14 Turbulence spectra for Re=200,000 case 

0.5 0.6 0.7 0.8 0.9 1

1

2

3

4

5

s/Ls

H

Re=25,000

Re=50,000

Re=100,000

Re=150,000

Re=200,000

Re=300,000

Fig. 15 Shape factor 

reduces the shape factor, indicating that separation is delayed. At the 

second station in Fig. 15, H is between 4 and 5 in all cases, indicating 

that the boundary layer is separated. Note that these high H values are 

not quantitatively accurate since the hot-wire is unable to accurately 

measure velocity within the separation bubble. The high H values, do, 

however, provide a good qualitative indicator that the flow is locally 

separated. By the third station, H has dropped to an attached turbulent 

boundary layer level for Re≥200,000 and by station 4 all but the two 

lowest Re cases are fully reattached. In the Re=25,000 case H remains 

NASA/CR—2012-217415 57
8 Copyright © 2008 by ASME 

200 400 600 800 1000 1200 1400

0

1

2

3

4

5

6

7

8

x 10−3

Cf=0.0256/Reθ

0.25

Cf=0.246⋅ 10−0.678H Reθ

−0.268

Reθ

Cf

Re=100,000

Re=150,000

Re=200,000

Re=300,000

Flat plate

Ludwieg−Tillmann

Fig. 16 Skin friction coefficient for high Re cases 

high, since the boundary layer never reattaches. At Re=50,000, H at 

the downstream stations lies between the attached values at high Re

and the separated values at Re=25,000, indicating that the boundary 

layer has reattached but has not fully recovered from the separation. 

The skin friction coefficient is shown as a function of momentum 

thickness Reynolds number in Fig. 16 for the high Re cases. It rises 

from zero after separation and then decreases in the streamwise 

direction. Also shown is a standard flat-plate correlation for turbulent 

boundary layers from Schlichting [29] and the Ludwieg-Tillmann 

correlation which accounts for pressure gradient effects. The present 

Cf values are as much as 50% lower than the flat-plate correlation and 

agree well with the Ludwieg-Tillmann correlation. 

Transition Correlations 

 Several correlations for predicting the starting location for separated 

flow transition are available in the literature, and some of these are 

tested below against the present data. The correlation equations are 

available in Volino [15]. Most of the correlations predict the distance 

from the separation point to the start of transition. Mayle [2] presents 

correlations for short and long separation bubbles which depend on 

Reθ at separation. Davis et al. [30] uses only the freestream turbulence 

intensity. Due to the acceleration through the passage, the local TI

dropped from the inlet value to about 1.3% at separation, and was set 

to this value for use with all correlations. Hatman and Wang [31] 

identify several transition modes and present correlations for each of 

them. Their laminar separation mode transition correlation depends on 

Res (based on local freestream velocity and distance from the leading 

edge to the separation point). Yaras [32] uses Res and a turbulence 

factor based on the freestream turbulence intensity and its integral 

length scale. Praisner and Clark [7] use Res and Reθ at separation. An 

additional correlation proposed by Roberts and Yaras [27] is 

Rest = (785-30 TF) Reθs

0.7 (3) 

where TF=TI(ss/λ)

0.2, ss is the distance from the leading edge to the 

separation point and λ is the integral length scale of the freestream 

turbulence. 

 While the above correlations are based on the conditions at the 

separation location, Volino and Bohl [33] reasoned that instabilities 

begin to grow when the boundary layer becomes unstable at the start 

of the adverse pressure gradient region. They compute the distance 

from the suction peak to the start of transition based on TI and Reθ at 

the suction peak. 

Table 3: Conditions at suction peak and separation location 

based on laminar boundary layer calculation 

Re sp/Ls ss/Ls Reθp Reθs

25,000 0.438 0.496 47 60 

50,000 0.438 0.536 64 89 

100,000 0.493 0.538 106 127 

150,000 0.493 0.539 130 157 

200,000 0.493 0.538 150 181 

300,000 0.493 0.538 183 221 

0 0.5 1 1.5 2 2.5 3 3.5

x 10

5

0.5

0.6

0.7

0.8

0.9

1

1.1

1.2

1.3

1.4

Re

s

t/Ls

Volino and Bohl [33]

Mayle [2] short bubble

Mayle [2] long bubble

Davis et al. [30]

Hatman and Wang [31]

Yaras [32]

Praisner and Clark [7]

Roberts and Yaras [27]

Present Experiment

Fig. 17 Predicted and experimental transition start location, 

error bars indicate uncertainty due to finite spacing of 

measurement stations 

 The location of the suction peak for each case was taken from the 

experimental data of Fig. 2. The separation location and the 

momentum thickness Reynolds numbers at separation, Reθs, and at the 

suction peak, Reθp, were estimated using a laminar calculation with the 

measured pressure gradient using the TEXSTAN [34] boundary layer 

code, as described in Volino [15]. The Reynolds numbers used with 

the correlations are listed in Table 3. The various correlations were 

then used to compute st, shown in Fig. 17, which is defined as the 

distance from the leading edge to the start of transition. The 

experimental transition locations, as determined from Figs. 7-14 are 

also shown. The uncertainties in the experimental locations, which 

result from the finite spacing between the measurement stations, are 

indicated by the error bars. Values of st/Ls>1 in the figure indicate that 

transition is predicted downstream of the trailing edge of the airfoil. 

 The Mayle [2] short bubble correlation, which is only intended for 

separation bubbles that reattach, predicts transition well in most cases, 

as does the Yaras [32] correlation. The Volino and Bohl [33] 

correlation agrees with the experiment at all Re. The Davis et al. [30] 

and Roberts and Yaras [27] correlations predict transition well for 

Re≥100,000, but too far downstream at lower Re. The other 

correlations predict transition too far downstream, but approach the 

experimental results at high Re. With low TI, all the correlations 

predicted the data well at high Re. At low Re, the Mayle [2] short 

bubble correlation predicted transition too far upstream, the Roberts 

and Yaras [27] and Volino and Bohl [33] correlations agreed with the 

data, and the other correlations predicted transition too far 

downstream. 

NASA/CR—2012-217415 58
9 Copyright © 2008 by ASME 

 (a) (b) (c) 

Fig. 18 Flow field on suction side of airfoil showing u′/Ue

(contours) and mean velocity (vectors): a) Re=25,000, b) 

100,000, c) 300,000 

Numerical Results 

 Velocity vector fields and contours of u′/Ue are shown in Fig. 18 for 

the simulation results using the Transition-sst model. As will be 

shown below, the Transition-sst model generally performed better than 

the SKW-sst model. The flow along the suction side of the airfoil is 

shown in the figure. For reference, the white lines in the figure 

correspond to the experimental measurement stations listed in Table 2. 

The short white line upstream of the others indicates the location of 

the suction peak. In agreement with the experimental results, the 

simulation predicts a burst separation bubble at Re=25,000. The 

turbulence level is high in the shear layer above the separation bubble, 

but the boundary layer does not reattach. The velocity vectors show a 

significant reduction in flow turning. At Re=100,000 and 300,000, 

again in agreement with the experiment, the boundary layer is 

attached. The turbulence contours indicate a thicker boundary layer at 

Re=100,000 than at 300,000. The location of the suction peaks and 

separation locations in Fig. 18 agree with those in Table 3 to within 

0.013 and 0.034 in sp/Ls and ss/Ls respectively. This lends support to 

the simple laminar boundary layer calculations used to produce the 

estimates in Table 3. The Reθp values which can be extracted from the 

simulation are about 11% lower than the values in Table 3. This 

difference is not large enough to significantly change the correlation 

predictions in Fig. 17. 

 To quantify the simulation results, pressure profiles for the 

numerical simulations are compared to the experimental results in Fig. 

19. At all Reynolds numbers, both the SKW-sst and Transition-sst 

models agree reasonably well with the experimental data. At 

Re=25,000, the Transition-sst model agrees with the data to within the 

experimental uncertainty from s/Ls=0.1 to 0.6. The SKW-sst Cp

prediction is slightly lower, but both models correctly predict that the 

boundary layer does not reattach. The Transition-sst model predicts a 

drop in Cp downstream of s/Ls=0.6 which is not seen in the data. The 

drop corresponds to the start of transition, as will be discussed below. 

At Re=100,000 (Fig. 18b), both models correctly predict an attached 

boundary layer. In the low TI case at this Re, both models correctly 

predicted a burst separation bubble, so they appear to handle the 

freestream turbulence effect correctly. The Transition-sst model 

provides a better prediction than the SKW-sst model at Re=100,000, 

particularly between s/Ls=0.5 and 0.7 where the data and the 

Transition-sst models show a slight plateau in Cp. The plateau 

indicates a boundary layer on the verge of separation or possibly a 

small separation bubble. For the Re=300,000 case (Fig. 18c), neither 

model or the experiment show any indication of separation. Both 

models agree well with the experiment, with the Transition-sst model 

providing a slightly better prediction downstream of the suction peak. 

The simulations under predict the peak Cp in the Re=100,000 and 

300,000 cases, and although the difference is not large, it is consistent 

0 0.2 0.4 0.6 0.8 1

0

0.5

1

1.5

2

2.5

s/Ls

Cp

(a) Exp.

SKW−sst

Trans−sst

0 0.2 0.4 0.6 0.8 1

0

0.5

1

1.5

2

2.5

s/Ls

Cp

(b)

0 0.2 0.4 0.6 0.8 1

0

0.5

1

1.5

2

2.5

s/Ls

Cp

(c)

Fig. 19 Cp profiles comparing simulations to experiment: 

a) Re=25,000, b) 100,000, c) 300,000 

with the difference between the data and the inviscid solution in Fig. 2. 

Bons et al. [14] saw a similar difference between their experimental 

data and simulations for similar conditions. 

 Mean velocity profiles for the Re=25,000 case are compared to the 

experiment in Fig. 20. Up to s/Ls=0.59, the turbulence model 

predictions agree well with each other and the experiment. 

Differences from the experiment inside the separation bubble are not 

meaningful, since the hot-wire cannot distinguish the direction of 

reverse flow. Downstream of s/Ls=0.59, the models correctly predict 

the growth of the separation bubble. The Transition-sst model 

generally provides a better match to the data, although the SKW-sst 

model is closer at s/Ls=0.88. A thicker separation bubble in the 

simulation than the experiment, as shown at s/Ls=0.97, is expected in 

cases with burst separation bubbles, particularly at downstream 

NASA/CR—2012-217415 59
10 Copyright © 2008 by ASME 

−0.5 0 0.5 1 1.5 0

0.05

0.1

0.15

0.2

0.25

0.3

s/Ls=0.53

y/Ls

Exp.

SKW−sst

Trans−sst

−0.5 0 0.5 1 1.5 0

0.05

0.1

0.15

0.2

0.25

0.3

s/Ls=0.59

−0.5 0 0.5 1 1.5 0

0.05

0.1

0.15

0.2

0.25

0.3

s/Ls=0.69

−0.5 0 0.5 1 1.5 0

0.05

0.1

0.15

0.2

0.25

s/Ls=0.78

y/Ls

U/Ue

−0.5 0 0.5 1 1.5 0

0.05

0.1

0.15

0.2

0.25

s/Ls=0.88

U/Ue

−0.5 0 0.5 1 1.5 0

0.05

0.1

0.15

0.2

0.25

s/Ls=0.97

U/Ue

Fig. 20: Comparison of simulation and experimental mean 

velocity profiles, Re=25,000 

0 0.1 0.2 0

0.05

0.1

0.15

0.2

0.25

0.3

s/Ls=0.53

y/Ls

Exp.

Trans−sst

0 0.1 0.2 0

0.05

0.1

0.15

0.2

0.25

0.3

s/Ls=0.59

0 0.1 0.2 0

0.05

0.1

0.15

0.2

0.25

0.3

s/Ls=0.69

0 0.1 0.2 0

0.05

0.1

0.15

0.2

0.25

s/Ls=0.78

y/Ls

u′/Ue

0 0.1 0.2 0

0.05

0.1

0.15

0.2

0.25

s/Ls=0.88

u′/Ue

0 0.1 0.2 0

0.05

0.1

0.15

0.2

0.25

s/Ls=0.97

u′/Ue

Fig. 21: Comparison of Transition-sst and experimental 

u′/Ue profiles, Re=25,000 

stations. As noted above, the tailboard in the experiment suppresses 

the separation bubble somewhat in cases without reattachment, 

particularly on the airfoils closest to the tailboard. The simulation, 

with its periodic boundary conditions, corresponds to an infinite 

cascade with no tailboard effects. 

 The turbulence kinetic energy provides some information about how 

the simulation computes transition. The rms fluctuating streamwise 

velocity, u′, is estimated from the simulations using the computed 

turbulence kinetic energy and an assumption of isotropic turbulence. 

This assumption is clearly not strictly correct, but allows an estimate 

of u′ to compare to the experimental data. This comparison is shown 

in Fig. 21 for the Transition-sst model in the Re=25,000 case. At 

s/Ls=0.53, there is a near wall peak in the experimental data which is 

not captured by the calculation. This is upstream of transition, so the 

peak is likely caused by the freestream turbulence buffeting the 

boundary layer. At all the other stations the peak is predicted at the 

correct distance from the wall. The simulation peak is still too low at 

s/Ls=0.59 and about 40% too high at s/Ls=0.69, but at all the other 

stations the simulation and experiment peak magnitudes match closely. 

An exact match should not be expected given the approximation in 

estimating u′ from the turbulence kinetic energy. Above the peak the 

match is also good, but closer to the wall the simulation shows higher 

fluctuating velocity than the experiment. Some of the difference may 

be attributable to the inability of the hot-wire to measure velocity 

accurately inside the separation bubble, but it appears that the 

simulation predicts a thicker shear layer than the experiment. Ibrahim 

et al. [16] also saw a thicker shear layer than the experiment in the low 

TI cases. 

 The rise in turbulence in the simulation can be used to predict the 

start of transition. The transition start location is designated as the 

location of the local maximum in turbulence kinetic energy following 

the suction peak, determined using the contours of Fig. 18. This 

particular location is a somewhat arbitrary choice, as one could also 

reasonably designate a location slightly upstream of the peak as the 

transition start. The local peak is, however, an unambiguous location 

and is useful for comparing transition among different cases. In the 

Re=25,000 case, the simulation predicts transition start at s/Ls=0.64, 

which agrees with the experimental location indicated in Figs. 8 and 

17. The start of transition corresponds to the start of the slow drop in 

Cp shown in Fig. 19a. The mixing associated with transition makes a 

shear layer more likely to reattach. Perhaps the transition and 

turbulence predicted by the Transition-sst model is pushing the shear 

layer closer to reattachment than observed in the experiment, which 

could cause the lower Cp. The thicker shear layer in the simulation 

may cause this effect. The effect is not large enough to cause a full 

reattachment in the simulation, so the simulation and experiment 

remain in overall good agreement. Ibrahim et al. [16] did not see the 

same drop Cp in their low TI Re=25,000 case, possibly because 

transition did not occur until s/Ls=0.74, and the separation bubble had 

become too thick for any hint of reattachment. In their Re=100,000 

case, Ibrahim et al. [16] saw transition start at s/Ls=0.59, and it 

induced the same slow drop in Cp shown in Fig. 19a. 

 Figures 22 and 23 show the mean velocity and u′ profiles for the 

Re=100,000 case. The Transition-sst model predicts the data well at 

most locations. Some difference in the shape of the mean profile is 

visible at the two most downstream stations, and the magnitude of the 

u′ peak is under predicted at the upstream stations. The freestream 

turbulence level in the experiment has also decayed more than in the 

simulation. The SKW-sst model does not do as well at the 

downstream stations, predicting a thicker boundary layer and a small 

separation bubble which were not observed in the experiment. 

Velocity and turbulence profiles for the Re=300,000 case are shown in 

Figs. 24 and 25. The Transition-sst model provides good agreement 

with the experimental mean profiles at all locations. With the same 

exceptions noted in Fig. 23, agreement is also good for the u′ profiles. 

The SKW-sst model again predicts a thicker boundary layer than the 

experiment at the downstream stations. The Transition-sst model 

predicts transition start at s/Ls=0.66 and 0.64 in the Re=100,000 and 

300,000 cases respectively. To within the experimental uncertainty, 

these locations agree with the experimental locations shown in Fig. 17. 

 Total pressure loss coefficients are shown in Fig. 26. The 

magnitude and width of the loss peaks in the experiments and 

simulations generally agree, showing the correct trend with Reynolds 

number. In the Re=25,000 case (Fig. 26a), the simulations predict 

higher loss peaks and less flow turning than the experiment. This is 

expected since the tailboard suppresses the separation bubble 

somewhat in the experiment, as noted above. A thicker bubble in the 

simulation will result in higher losses and divert the flow toward the 

pressure side of the passage, moving the loss peak to lower φ/Lφ. 

NASA/CR—2012-217415 60
11 Copyright © 2008 by ASME 

−0.5 0 0.5 1 1.5 0

0.05

0.1

0.15

0.2

0.25

0.3

s/Ls=0.53

y/Ls

Exp.

SKW−sst

Trans−sst

−0.5 0 0.5 1 1.5 0

0.05

0.1

0.15

0.2

0.25

0.3

s/Ls=0.59

−0.5 0 0.5 1 1.5 0

0.05

0.1

0.15

0.2

0.25

0.3

s/Ls=0.69

−0.5 0 0.5 1 1.5 0

0.05

0.1

0.15

0.2

0.25

s/Ls=0.78

y/Ls

U/Ue

−0.5 0 0.5 1 1.5 0

0.05

0.1

0.15

0.2

0.25

s/Ls=0.88

U/Ue

−0.5 0 0.5 1 1.5 0

0.05

0.1

0.15

0.2

0.25

s/Ls=0.97

U/Ue

Fig. 22: Comparison of simulation and experimental mean 

velocity profiles, Re=100,000 

0 0.1 0.2 0

0.05

0.1

0.15

0.2

0.25

0.3

s/Ls=0.53

y/Ls

Exp.

Trans−sst

0 0.1 0.2 0

0.05

0.1

0.15

0.2

0.25

0.3

s/Ls=0.59

0 0.1 0.2 0

0.05

0.1

0.15

0.2

0.25

0.3

s/Ls=0.69

0 0.1 0.2 0

0.05

0.1

0.15

0.2

0.25

s/Ls=0.78

y/Ls

u′/Ue

0 0.1 0.2 0

0.05

0.1

0.15

0.2

0.25

s/Ls=0.88

u′/Ue

0 0.1 0.2 0

0.05

0.1

0.15

0.2

0.25

s/Ls=0.97

u′/Ue

Fig. 23: Comparison of Transition-sst and experimental 

u′/Ue profiles, Re=100,000 

Between wakes the Transition-sst model indicates lower loss than the 

experiment. This difference may be related to the under prediction of 

the freestream turbulence decay noted in Figs. 23 and 25. The 

Re=100,000 and 300,000 cases (Figs. 26b and 26c) do not suffer from 

tailboard effects in the experiment since the boundary layer reattaches. 

In both cases the Transition-sst model predicts the peak magnitudes to 

within about 5%, but the peak location is shifted slightly toward the 

pressure side of the passage in the simulation. The SKW-sst model 

does not do as well, predicting a higher peak and more of a shift 

toward the pressure side. As in the lower Re case, both models under 

predict the loss between wakes. 

CONCLUSIONS 

 The flow over the very high lift L1A airfoil was studied under high 

freestream turbulence conditions. Reynolds numbers based on suction 

surface length and nominal exit velocity ranged from 25,000 to 

300,000. At the lowest Reynolds number, the laminar suction surface 

−0.5 0 0.5 1 1.5 0

0.05

0.1

0.15

0.2

0.25

0.3

s/Ls=0.53

y/Ls

Exp.

SKW−sst

Trans−sst

−0.5 0 0.5 1 1.5 0

0.05

0.1

0.15

0.2

0.25

0.3

s/Ls=0.59

−0.5 0 0.5 1 1.5 0

0.05

0.1

0.15

0.2

0.25

0.3

s/Ls=0.69

−0.5 0 0.5 1 1.5 0

0.05

0.1

0.15

0.2

0.25

s/Ls=0.78

y/Ls

U/Ue

−0.5 0 0.5 1 1.5 0

0.05

0.1

0.15

0.2

0.25

s/Ls=0.88

U/Ue

−0.5 0 0.5 1 1.5 0

0.05

0.1

0.15

0.2

0.25

s/Ls=0.97

U/Ue

Fig. 24: Comparison of simulation and experimental mean 

velocity profiles, Re=300,000 

0 0.1 0.2 0

0.05

0.1

0.15

0.2

0.25

0.3

s/Ls=0.53

y/Ls

Exp.

Trans−sst

0 0.1 0.2 0

0.05

0.1

0.15

0.2

0.25

0.3

s/Ls=0.59

0 0.1 0.2 0

0.05

0.1

0.15

0.2

0.25

0.3

s/Ls=0.69

0 0.1 0.2 0

0.05

0.1

0.15

0.2

0.25

s/Ls=0.78

y/Ls

u′/Ue

0 0.1 0.2 0

0.05

0.1

0.15

0.2

0.25

s/Ls=0.88

u′/Ue

0 0.1 0.2 0

0.05

0.1

0.15

0.2

0.25

s/Ls=0.97

u′/Ue

Fig. 25: Comparison of Transition-sst and experimental 

u′/Ue profiles, Re=300,000 

boundary layer separated and did not reattach. At the higher Reynolds 

numbers, a separation bubble was followed by transition and 

reattachment. Spectral peaks suggest that transition is caused by the 

growth of an instability, as opposed to a pure bypass, in spite of the 

high freestream turbulence. The transition locations in both high and 

low TI cases were well predicted by the correlation of Volino and Bohl 

[33]. The separation bubble became smaller as Reynolds number 

increased, and was effectively eliminated at the highest Reynolds 

numbers. The tendency toward separation at intermediate Reynolds 

numbers was still large enough to increase the boundary layer 

thickness and significantly increase pressure losses above the high Re

cases. The present results contrast with low freestream turbulence 

results acquired in the same facility. With low TI, the boundary layer 

did not reattach at intermediate Reynolds numbers, in spite of 

transition taking place in the separated shear layer. High freestream 

turbulence appears to increase the thickness of the shear layer over the 

separation bubble, thereby decreasing the bubble thickness. The 

NASA/CR—2012-217415 61
12 Copyright © 2008 by ASME 

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

(a)

φ/Lφ

ψ

B3

B4 B5

B6

Exp.

SKW−sst

Trans−sst

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

(b)

φ/Lφ

ψ

B3 B4 B5

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

(c)

φ/Lφ

ψ

B3 B4 B5

Fig. 26 Total pressure loss coefficient at 0.63 Cx

downstream of cascade comparing simulations to 

experiment: a) Re=25,000, b) 100,000, c) 300,000 

thinner bubble results in turbulence closer to the wall when transition 

begins, which promotes reattachment. Reattachment changes the 

pressure distribution on the airfoil, causing the suction peak to rise and 

move downstream and delaying separation. The result is higher lift 

and an even smaller separation bubble. 

 Numerical simulations with the 4 equation Transition-sst model of 

Menter et al. [21] correctly predicted separation, transition and 

reattachment at all Reynolds numbers. Some discrepancies between 

the model prediction and the experimental data were noted, but in 

general the model predicted well the pressure distribution on the 

airfoil, the total pressure losses, and mean and fluctuating velocity 

profiles along the suction surface of the airfoil. The k-ω-sst (SKW-sst) 

model of Menter [19] did not do quite as well. Comparing to the 

results of Ibrahim et al. [16], the simulations correctly predicted the 

major differences between the high and low freestream turbulence 

cases. 

ACKNOWLEDGMENTS

 This work was sponsored by the National Aeronautics and Space 

Administration. The grant monitor is Dr. Anthony Strasizar of the 

NASA Glenn Research Center. The support of the United States 

Naval Academy Technical Support Department Shop and Fluids 

Laboratory is greatly appreciated. 

 

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NASA/CR—2012-217415 62
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NASA/CR—2012-217415 63
1 

This material is declared a work of the U.S. Government and is not subject to copyright protection in the United States. Approved for public release; 

distribution is unlimited. 

Proceedings of ASME Turbo Expo 2009: Power for Land, Sea and Air 

GT2009 

June 8-12, 2009, Orlando, Florida, USA

GT2009-59983 

EXPERIMENTAL AND COMPUTATIONAL INVESTIGATIONS OF LOW-PRESSURE 

TURBINE SEPARATION CONTROL USING VORTEX GENERATOR JETS

Ralph J. Volino

Mechanical Engineering Department 

United States Naval Academy 

Annapolis, Maryland 21402-5042 

volino@usna.edu 

Olga Kartuzova and Mounir B. Ibrahim 

Mechanical Engineering Department 

Cleveland State University 

Cleveland, Ohio 44115-2425 

m.ibrahim@csuohio.edu 

ABSTRACT 

 Boundary layer separation control has been studied using vortex 

generator jets (VGJs) on a very high lift, low-pressure turbine airfoil. 

Experiments were done under low freestream turbulence conditions on 

a linear cascade in a low speed wind tunnel. Pressure surveys on the 

airfoil surface and downstream total pressure loss surveys were 

documented. Cases were considered at Reynolds numbers (based on 

the suction surface length and the nominal exit velocity from the 

cascade) of 25,000, 50,000 and 100,000. Jet pulsing frequency, duty 

cycle, and blowing ratio were all varied. In all cases without flow 

control, the boundary layer separated and did not reattach. With the 

VGJs, separation control was possible even at the lowest Reynolds 

number. Pulsed VGJs were more effective than steady jets. At 

sufficiently high pulsing frequencies, separation control was possible 

even with low jet velocities and low duty cycles. At lower 

frequencies, higher jet velocity was required, particularly at low 

Reynolds numbers. Effective separation control resulted in an 

increase in lift of up to 20% and a reduction in total pressure losses of 

up to 70%. Simulations of the flow using an unsteady RANS code 

with the four equation Transition-sst model produced good agreement 

with experiments in cases without flow control, correctly predicting 

separation, transition and reattachment. In cases with VGJs, however, 

the CFD did not predict the reattachment observed in the experiments. 

NOMENCLATURE 

B blowing ratio, maximum jet velocity/local freestream velocity 

Cp ( ) 2

T − ρU/PP2 e , pressure coefficient 

Cpint pressure coefficient integrated in axial direction 

Cpinv Cpint for invisid flow through cascade 

Cx axial chord length 

D duty cycle, fraction of time valve is open 

F fLj-te/Uave, dimensionless frequency 

f frequency 

Lj-te distance from VGJs to trailing edge

Ls suction surface length 

Lφ blade spacing (pitch) 

P pressure 

PS upstream static pressure 

PT upstream stagnation pressure 

PTe downstream stagnation pressure 

Re UeLs/ν, exit Reynolds number 

s streamwise coordinate, distance from leading edge 

T period of jet pulsing cycle 

t time 

Uave average freestream velocity between VGJs and trailing edge 

Ue nominal exit freestream velocity, based on inviscid solution 

x axial distance from leading edge 

φ coordinate along blade spacing, normal to axial chord 

ν kinematic viscosity 

ρ density 

ψ (PT-PTe)/(PT-PS), total pressure loss coefficient 

ψint total pressure loss integrated over blade spacing 

INTRODUCTION 

 Boundary layer separation can lead to partial loss of lift and higher 

aerodynamic losses on low-pressure turbine (LPT) airfoils (e.g. 

Hourmouziadis [1], Mayle [2], and Sharma et al. [3]). As designers 

impose higher loading to improve efficiency and lower cost, the 

associated strong adverse pressure gradients on the suction side of the 

airfoil can exacerbate separation problems. The problem is 

particularly relevant in aircraft engines at cruise conditions, due to the 

lower density and therefore lower Reynolds numbers at altitude. A 

component efficiency drop of 2% may occur between takeoff and 

cruise in large commercial transport engines, and the difference could 

be as large as 7% in smaller engines operating at higher altitudes [4, 

5]. Prediction and control of suction side separation, without sacrifice 

of the benefits of higher loading, is therefore, crucial for improved 

engine design. 

 Separation and separated flow transition, which can lead to 

boundary layer reattachment, have received considerable attention. 

Lists of various studies are provided by Volino [6, 7]. In general, 

previous work shows that the strong acceleration on the leading 

section of the airfoil keeps the boundary layer thin and laminar, even 

in the presence of elevated freestream turbulence. When separation 

does occur, it is usually just downstream of the suction peak. If 

transition then occurs in the shear layer over the separation bubble, it 

Proceedings of ASME Turbo Expo 2009: Power for Land, Sea and Air 

GT2009 

June 8-12, 2009, Orlando, Florida, USA

GT2009-59983

NASA/CR—2012-217415 64

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distribution is unlimited. 

is typically rapid and often causes the boundary layer to reattach [6, 8]. 

Bons et al. [9] and Volino [6], found that with very high loading, 

however, the separation bubble can become so thick that transition 

alone is not sufficient to cause reattachment. 

 One way to improve performance is to design airfoils with pressure 

gradients more resistant to separation, as described by Praisner and 

Clark [10]. Forward loading, for example, makes airfoils more 

separation resistant by extending the adverse pressure gradient on the 

aft portion of the suction side over a longer distance. This reduces the 

local pressure gradient at all locations, making separation less likely. 

If separation does occur, forward loading provides a longer distance 

along the airfoil surface for reattachment. Forward loading has some 

disadvantages, however. As noted by Zhang et al. [11], the longer 

region of turbulent flow on a forward loaded airfoil can lead to 

increased profile losses. Forward loading also creates longer regions 

of strong pressure gradient on the endwalls, which can produce 

stronger secondary flows and losses. If flow control were incorporated 

in the design of an advanced airfoil, as discussed by Bons et al. [12], it 

might be possible to produce an aft loaded airfoil that was resistant to 

separation and had low profile and secondary loss characteristics over 

a range of Reynolds numbers. 

 Separation control with passive devices such as boundary layer trips 

has been shown effective by Zhang et al. [11], Bohl and Volino [13], 

Volino [14], and others. Passive devices have the distinct advantage 

of simplicity, but they also introduce parasitic losses. Devices which 

are large enough to control separation at the lowest Reynolds numbers 

in an engine’s operating range would tend to produce higher than 

necessary losses at higher Reynolds numbers. Active devices could 

potentially provide better control over the entire operating range of 

interest and be reduced in strength or turned off to avoid unnecessary 

losses when they are not needed. 

 The literature contains many examples of active separation control. 

A few which could be applied in turbomachinery are discussed in 

Volino [15]. Plasma devices, as used by Huang et al. [16], could be 

viable, and are under active study. Vortex generator jets (VGJs), as 

introduced by Johnston and Nishi [17], are another alternative, and are 

the subject of the present study. Blowing from small, compound 

angled holes is used to create streamwise vortices. The vortices bring 

high momentum fluid into the near wall region, which can help to 

control separation. The most effective VGJs enter the boundary layer 

at a relatively shallow pitch angle (typically 30 to 45 degrees) relative 

to the wall and a high skew angle (45 to 90 degrees) relative to the 

main flow. Additionally, the jets promote transition, and turbulent 

mixing also helps to mitigate separation. Bons et al. [12] noted that in 

the case of pulsed VGJs, the turbulence effect was more significant 

than the action of the vortices. Bons et al. [4, 18], Volino [19], Volino 

and Bohl [20], McQuilling and Jacob [21], and Eldredge and Bons 

[22] all used VGJs on the highly loaded Pack B LPT airfoil. 

Separation was essentially eliminated, even at the lowest Reynolds 

number considered, (Re=25,000 based on suction surface length and 

nominal exit velocity). Pulsed jets were more effective than steady 

jets. The initial disturbance created by each pulse caused the boundary 

layer to attach. The turbulence was followed by a calmed period 

(Gostelow et al. [23] and Schulte and Hodson [24]) during which the 

boundary layer was very resistant to separation, much like a turbulent 

boundary layer, but very laminar-like in terms of its fluctuation levels 

and low losses. When the time between pulses was long enough, the 

boundary layer did eventually relax to a separated state, but due to the 

control which persisted during the calmed period, the VGJs were 

effective even with low jet pulsing frequencies, duty cycles and mass 

flow rates. Since the boundary layer was attached and undisturbed for 

much of the jet pulsing cycle, profile losses were low. 

 Similar results with pulsed VGJs were found on the L1M airfoil by 

Bons et al. [25]. The L1M is more highly loaded than the Pack B, but 

more resistant to separation because of forward loading. A large 

separation bubble followed by boundary layer reattachment was 

observed at low Reynolds numbers, and pulsed VGJs reduced the size 

of the bubble. 

 In the present study, the very highly loaded L1A airfoil is used. The 

L1A was designed at the Air Force Research Laboratory (AFRL) and 

is available on a limited basis from Clark [26]. It is an aft loaded 

blade with the same flow angles and loading as the L1M. Dimensions 

of the L1A as used in the present study are given in Table 1. Based on 

the design calculations of Clark [26], the L1A has a Zweifel 

coefficient of 1.35, which corresponds to 10% higher loading than the 

“ultra-high lift” airfoils described by Zhang and Hodson [27], and 17% 

higher loading than the Pack B. Because the L1A is aft loaded, it is 

more prone to separation than the L1M, as documented in Bons et al. 

[9], Volino [6], Ibrahim et al. [28], and Volino et al. [29]. In cases 

without flow control and with low freestream turbulence, the boundary 

layer separates when Re<150,000 and does not reattach, in spite of 

transition to turbulence in the shear layer over the separation bubble in 

all cases. This result contrasts with the results of earlier studies on less 

aggressive airfoils, which all showed reattachment after transition. 

The separation bubble on the L1A is about four times thicker than that 

on the Pack B. The larger distance from the shear layer to the wall on 

the L1A apparently prevents the turbulent mixing in the shear layer 

from reaching the wall and causing reattachment. The failure of the 

boundary layer to reattach results in a 20% loss in lift and increases 

profile losses by a factor of 7. At higher Reynolds numbers the 

separation bubble is small and the boundary layer is attached over 

most of the airfoil. In cases with high freestream turbulence, results 

are similar, but the shear layer is somewhat thicker and the separation 

bubble thinner due to increased mixing induced in the shear layer. 

This results in reattachment after transition at Re=50,000 and 100,000. 

At the lowest Re considered (25,000) the boundary layer still does not 

reattach. 

 Attempts to simulate LPT flows computationally have met with 

varying degrees of success. A few are discussed in Ibrahim et al. [28]. 

Most turbulence models, such as the standard k-ε model, fail to predict 

separation, typically because transition is predicted too far upstream. 

A few have more success in cases without flow control. Suzen et al. 

[30], for example, successfully predicted the flow over the Pack B 

airfoil for a wide range of conditions using an intermittency transport 

model. Menter et al. [31] and Langtry et al. [32] used an intermittency 

transport equation to trigger transition onset along with a transport 

equation for the transition momentum thickness Reynolds number. 

These two equations were coupled with a shear stress transport 

turbulence model (SST). The combined model, referred to here as the 

Transition-sst model, successfully predicted separation, transition, and 

reattachment on the Pack B airfoil. Gross and Fasel [33] used course 

grid direct numerical simulations (DNS), implicit large eddy 

simulations (ILES) and unsteady Reynolds averaged Navier-Stokes 

(URANS) models to predict Pack B flows. The DNS and ILES results 

agreed when the computational grid was sufficiently fine, and some of 

the URANS models agreed as well. Agreement with experimental 

data was good in some instances, but significant differences were 

observed in others. This was attributed to possible differences 

between the inlet flow conditions in the experiment and computations. 

Rizzetta and Visbal [34] used DNS and found reasonable agreement 

with Pack B experiments, although differences in velocity profiles 

were observed at some locations. 

 The flow over the L1A airfoil was computed by Ibrahim et al. [28] 

and Volino et al. [29] using the Transition-sst model. As with the 

NASA/CR—2012-217415 65
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distribution is unlimited. 

Table 1: Cascade parameters 

Axial 

Chord, Cx

[mm] 

True 

Chord 

[mm] 

Pitch, 

Lφ

[mm] 

Span 

[mm] 

Suction 

side, Ls

[mm] 

Inlet 

flow 

angle 

Exit 

flow 

angle 

134 146 136 724 203 35° 60° 

Pack B airfoil, this model correctly predicted separation, transition and 

reattachment for both high and low freestream turbulence and at high 

and low Reynolds numbers. Agreement with experimental total 

pressure loss surveys and velocity profiles in the suction side boundary 

layer were also good. 

 A few attempts have been made to simulate LPT flows with vortex 

generator jets. Garg [35] used a RANS calculation utilizing the k-ωsst model of Menter [36]. Although some trends were correct, the 

flow control seen in the experiments was not well predicted. Rizzetta 

and Visbal [34] had more success with their DNS, but differences 

from experimental results were still observed. 

 In the present study, flow control with vortex generator jets is 

investigated experimentally and computationally on the L1A airfoil. 

As noted above, the L1A provides a good test case for flow control 

since it is a modern aggressive airfoil with severe separation problems 

at low Reynolds numbers. Bons et al. [9] have shown that VGJs can 

be effective with the L1A for some cases. The present study considers 

a wide range of cases with different jet pulsing frequencies, amplitudes 

and duty cycles. Low (0.6%) freestream turbulence cases are 

considered with Reynolds numbers ranging from 25,000 to 100,000, as 

these were the cases with significant separation bubbles in the baseline 

experiments. For the computations URANS is used since it could 

provide a less computationally expensive design tool than LES or 

DNS. The Transition-sst model is used, as it was the most successful 

of the models used in the baseline cases. 

EXPERIMENTAL FACILITY AND MEASUREMENTS

 Experiments were conducted in a closed loop wind tunnel with a 

seven blade linear cascade located in the wind tunnel’s third turn, as 

shown in Fig. 1a. A fine screen located upstream of the cascade is 

used to break up the boundary layers which form upstream of the test 

section and to provide uniform inlet conditions to the cascade. The 

freestream turbulence entering the cascade was measured with a crosswire probe positioned just upstream of the center blade. The 

turbulence intensity is 0.8% in the streamwise component and 0.5% in 

the cross stream components. The integral length scale of the 

streamwise component is 6.3 cm. Further details of the facility and 

inlet flow are in Volino [6]. The freestream turbulence intensity in an 

engine is expected to be of the order 4%, which is significantly higher 

than in the present experiments. The freestream length scales are 

expected to be smaller in the engine. Higher intensity and smaller 

length scales promote more rapid transition of the boundary layer and 

more resistance to separation, as shown in Volino et al. [29]. The 

effect of wakes from upstream airfoils would be similar. The present 

experiments provide a lower bound for the expected freestream 

conditions in an engine and a more challenging case for flow control. 

The present results will serve as a baseline for future work with higher 

freestream turbulence and wakes, allowing separation of the effects of 

flow control, freestream turbulence and wakes. Better understanding 

of the separate effects may lead to better prediction of the combined 

effects under engine conditions. 

 A tailboard, shown in Fig. 1a, is needed to produce the correct exit 

flow angle from the cascade. Its position was set to produce 

periodicity at high Reynolds numbers. A tailboard on the opposite 

side of the cascade, and inlet guide vanes were found to be 

unnecessary. To produce the correct approach flow to the end blades 

tailboard

flap

flap

screen

B1

B2

B3

B4

B5

B6

B7

a) 

(b) 

Fig. 1 Drawings of test section: a) linear cascade, b) airfoil 

with VGJ holes and cross section of hole geometry 

(B1 and B7), the amount of flow escaping around the two ends of the 

cascade is controlled with the flaps shown in Fig. 1a. The inlet flow 

angle was checked with a three-hole pressure probe and found to agree 

with the design angle to within 2° of uncertainty. Good periodicity at 

high Reynolds numbers was obtained in the exit flow. At low 

Reynolds numbers, when significant separation bubbles are present, 

the periodicity is not as good due to suppression of the separation 

bubble thickness on the blades closest to the tailboard. In cases with 

effective flow control, periodicity is reestablished. The lack of 

periodicity in cases with large separation bubbles is considered 

acceptable since the focus of the study is separation control, and not 

documentation of cases with large separation that would be 

unacceptable in practice. This compromise facilitates the study of a 

larger number of cases with flow control by obviating the need to 

adjust the tailboard by trial and error for each case. It also provides for 

better repeatability in the experiments, since the position of the 

tailboard is fixed for all cases. Any improvements made with flow 

control will be larger in practice than documented in the experiment, 

due to the effect of the tailboard in suppressing the bubble size in the 

uncontrolled cases. 

 Each blade in the cascade has a central cavity which extends along 

the entire span. The cavity is closed at one end and has a fitting at the 

opposite end connected to a compressed air line. Air is supplied to the 

cavities from a common manifold. Manual ball valves are placed in 

the tubing between the manifold and blades to insure that each blade 

receives the same air flow. The valves also help to damp high 

frequency oscillations in the jet velocity when the VGJs are pulsed. 

The manifold is supplied through two fast response solenoid valves 

(Parker Hannifin 009-0339-900 with General Valve Iota One pulse 

NASA/CR—2012-217415 66
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distribution is unlimited. 

driver) operating in parallel. The valves are supplied through a 

pressure regulator by the building air supply. A single spanwise row 

of holes was drilled into the suction surface of each blade at the 

inviscid pressure minimum location, s/Ls=0.5 (x/Cx=0.62), where s is 

the distance from the leading edge and Ls is the suction surface length. 

The pressure minimum has been shown in the studies listed above to 

be about the optimal location for flow control devices. The effects of 

devices located farther upstream are damped by the favorable pressure 

gradient, and devices located downstream of the separation point can 

also lose effectiveness. The holes are 0.8 mm in diameter and drilled 

at 30° to the surface and 90° to the main flow direction, as shown in 

Fig. 1b. This is the same orientation used in all the VGJ studies listed 

above. The hole spacing is 10.6 diameters, and the length to diameter 

ratio is 12. 

 The solenoid valves pulse the VGJs, and the pulsing frequency is 

presented below in dimensionless form as F=fLj-te/Uave, where Lj-te is 

the streamwise distance from the VGJ holes to the trailing edge, and 

Uave is the average freestream velocity over this distance. For flow 

over single airfoils, F≥1 is typically needed to maintain separation 

control, but for cascades, Bons et al. [18] showed that control is 

possible in some cases with F=0.1. As shown in Volino [15] and Bons 

et al. [18], this is due to the extended calmed region which follows the 

jet disturbance. In practice, VGJs could be timed to wake passing in 

an LPT, which has a typical frequency of about F=0.3. 

 The center blade, designated B4 in Fig. 1, contains pressure taps 

near the spanwise centerline. Pressure surveys are made using a 

pressure transducer (0-870 Pa range Validyne transducer). Stagnation 

pressure is measured with a pitot tube upstream of the cascade. The 

uncertainty in the suction side pressure coefficients is 0.07. Most of 

this uncertainty is due to bias error. Stochastic error is minimized by 

averaging pressure transducer readings over a 10 second period. 

 Total pressure losses are documented using a Kiel probe traversed 

across three blade spacings, 0.63 Cx downstream of the cascade. A 

traverse is located in the wind tunnel downstream of the cascade to 

move the probe. The traverse causes an acceptably low blockage 

when it is located at least two Cx downstream of the cascade. 

 Pressure and loss surveys were acquired at nominal Re=25,000, 

50,000, 100,000. The Reynolds number, as defined above, is based on 

the suction surface length and the nominal cascade exit velocity. The 

corresponding Reynolds numbers based on the cascade inlet velocity 

and the axial chord length are 10,000, 20,000 and 40,000. 

NUMERICAL SIMULATIONS 

 Calculations were done using the commercial code Fluent with a 

URANS solver and the Transition-sst model of Menter et al. [31]. The 

three dimensional computational domain includes a single passage. 

The boundary conditions on the sides of the passage are periodic. The 

exit boundary is located 3.8 Cx downstream of the trailing edges in the 

flow direction. Zero gage pressure is specified at the exit. As 

described in Ibrahim et al. [28], exit locations extending farther 

downstream were tested to insure the results were independent of the 

location chosen. A uniform velocity inflow condition is specified 1.9 

Cx upstream of the blade leading edges in the flow direction. The inlet 

flow angle is set to 33°

 based on an inviscid calculation of the full 

cascade shown in Fig. 1. This angle agrees with the experimentally 

measured inlet angle to within the experimental uncertainty. In the 

spanwise direction, the domain includes one VGJ, and periodic 

boundary conditions are used. The full length of the hole is included 

in the simulations, allowing the jet velocity profile to develop before 

entering the main flow. A uniform velocity boundary condition is 

specified at the hole inlet. For the pulsed jet cases, the inlet velocity is 

set as a square wave. The upstream plenum is not included in the 

calculations. Grid independent results were obtained using 1.5 million 

cells with 593 nodes on the airfoil (362 nodes on the suction surface 

and 230 nodes on the pressure surface). The closest grid points to the 

wall were within one viscous unit (y

+<1) on all surfaces. A structured 

grid is clustered around the VGJ exit, and an unstructured grid is used 

for the rest of the computational domain. Boundary layer grid 

resolution is used inside of the jet tube. The three dimensional grid 

was checked by comparing results for a baseline case to previous 

results with a two dimensional grid. The periodic boundary conditions 

were checked by performing calculations for a two channel domain 

and a domain with three VGJs in the spanwise direction. 

RESULTS 

Jet Velocity 

 The center blade in the cascade includes a pressure tap in the central 

cavity to measure the supply pressure to the VGJs. The jet velocity is 

calibrated against the cavity pressure using a hot-film probe with a 

0.25 mm active sensor length located at the center of the exit plane of 

a VGJ hole. Calibration is done with the main flow in the wind tunnel 

turned off. In subsequent experiments, the desired jet velocity is 

produced by setting the difference between the cavity pressure and the 

static pressure in the freestream. During calibration, for each case of 

interest, the time averaged pressure in the cavity is measured along 

with the instantaneous jet velocity. For pulsed jet cases, the velocity is 

then phase averaged over the jet pulsing cycle. A typical result is 

shown in Fig. 2. The blowing ratio, B, is defined as the ratio of the 

measured jet velocity to the expected freestream velocity at the jet 

exit, based on an inviscid flow solution for the cascade. In the figure, 

the time during the cycle, t, is normalized on the pulsing period, T. 

The examples shown in Fig. 2 correspond to a pulsing frequency of 3 

Hz, with duty cycles, D, of 2%, 10%, and 50%. The duty cycle is the 

fraction of time the solenoid valve is open during each pulsing cycle. 

The nominal blowing ratio is set as the maximum value measured 

during the cycle. The maximum typically occurs shortly after the 

solenoid valve is opened, and is followed by a decay toward a steady 

value. The maximum is used because the initial pulse from the jet is 

believed to be most influential for flow control. Since the calibration 

is done with the main flow off, the freestream velocity used to define 

the blowing ratio (B=1) for Fig. 2 is arbitrary, and was chosen to 

corresponds to the lowest Reynolds number case (Re=25,000) 

considered in the study. For a Re=50,000 case, the results in Fig. 2 

would correspond to B=0.5. As shown in Fig. 2, the velocity from the 

jets does not form a perfect square wave when the jets are pulsed. The 

decay after the initial peak and other oscillations are believed to result 

from viscous losses and pressure waves reflecting in the manifold, 

tubing and cavity upstream of the VGJ holes. Similar behavior was 

observed by Bons et al. [9]. The oscillations shown in Fig. 2 are 

repeatable, as shown by the agreement between the three cases shown 

during the valve-open part of the cycle. The exact pattern of the 

oscillations is expected to be facility dependent. The success of the 

VGJs at low duty cycle, as shown below, suggests that it is the initial 

velocity peak, and not the oscillations which follow, which are most 

important for flow control. Further testing in a different facility or 

through CFD would be useful to confirm this. 

 It should be noted that the jet velocity used to define the blowing 

ratio is the measured value from the hot film probe. The jets are 

expected to be laminar with nearly parabolic velocity profiles, at least 

under steady conditions, since the length to diameter ratio of the holes 

is large. This is confirmed by the CFD predictions shown in Fig. 3. 

The maximum velocity in each profile is about 1.8 times the mean, 

which is near the fully developed value of 2 for a parabolic profile. 

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Fig. 2 Ensemble averaged jet velocity at f=3 Hz, with D=2%, 

10% and 50% 

2.73375, m/s

4.84461, m/s

8.8008, m/s

0.00

1.00

2.00

3.00

4.00

5.00

6.00

7.00

8.00

9.00

0 0.2 0.4 0.6 0.8 1

(distance across hole)/D

Jet exit velocity, m/s

Jet inlet vel.= 1.573 m/s (B=1)

Jet inlet vel.= 2.69 m/s (B=1.8)

Jet inlet vel.= 4.719 m/s (B=3)

Fig. 3 CFD prediction of VGJ velocity profile at hole exit 

Given a parabolic profile and the length of the hot film sensor, the 

average velocity over the sensor is 0.97 times the maximum velocity at 

the jet center. The measured velocity, therefore, essentially matches 

the maximum jet velocity. 

 Examples of CFD results for pulsed film cooling cases are shown in 

Fig. 4 for cases with 12 Hz pulsing. With D=10%, the CFD provides a 

good approximation to the experiment. The ratio of the maximum jet 

exit velocity to inlet velocity is about 1.7. Computed results with 

longer duration pulses did not show the expected decay after the initial 

maximum, indicating a difference in the imposed velocity at the hole 

inlet between CFD and experiment. Computational results presented 

below are limited to cases with f=12 Hz and D=10%. 

Re=25,000 

 Pressure profiles for all cases with Re=25,000 are shown in Fig. 5. 

The inviscid profile for the L1A airfoil is shown for comparison. 

Figure 5a shows that with B=0.25, there is little change from the 

baseline case. The VGJs are ineffective. With B=0.5 in Fig. 5b, there 

is no evidence of control at the two lowest frequencies, but at F=0.56 

(12 Hz) and 1.12, there is a drop in Cp near the trailing edge, 

indicating at least partial reattachment. This drop occurs even with a 

duty cycle of 10%. A lower duty cycle at high frequency could not be 

achieved in the experiments due to limitations in the solenoid valve 

response time. Increasing to B=0.75, Fig. 5c shows a small effect of 

pulsing at F=0.28, particularly with D=50%, and a stronger effect at 

the higher frequencies independent of duty cycle. Although the 

boundary layer appears to reattach near the trailing edge at the higher 

frequencies, a plateau in Cp up to s/Ls=0.7 indicates a large separation 

bubble is still present. As the blowing rate is increased to B=1.0, 1.5, 

Fig. 4 Comparison of measured and computed VGJ exit 

velocity with inlet velocity also shown for CFD results 

2.0, 2.5 and 3.0 in Figs. 5d-5h, there is little or no control at F=0.14 

unless B is larger than 2.0 and D=50%. At F=0.28, the control gets 

progressively better as B increases and tends to be better at the higher 

duty cycles. For the higher frequencies, B=0.75 appears to be 

sufficient, and raising B further does not appear to improve the results. 

A high blowing ratio of 3.0 might be impractical for application if the 

jets choke before reaching this velocity. If achievable, however, the 

VGJ mass flow rate with steady blowing at B=3.0 would be only about 

0.2% of the main flow. The mass flow rate with lower B or pulsing 

would be even lower. 

 The loss coefficient, ψ, is shown in Fig. 6. The coordinate φ

indicates the distance in the direction perpendicular to the axial chord. 

The normalizing quantity Lφ is the blade spacing. The origin, φ=0, 

corresponds to the location directly downstream of the trailing edge of 

the center blade in the direction of the exit design flow angle. Low 

blowing ratio cases are shown in Fig. 6a. Only the higher frequencies 

were considered at low B since the Cp results indicate that lower 

frequency pulsing is ineffective. The pulsed VGJs help to reduce the 

magnitude of the loss peaks by about 20%. The peaks are still much 

higher and wider (by about a factor of 2) than those observed in the 

high Re≥200,000 baseline cases [6]. The peaks are also shifted to the 

left of the design positions of φ/Lφ =-1, 0 and 1, indicating that 

compared to the high Re cases there is a 6 degree reduction in flow 

turning caused by the separation bubble. This is still better than the 13 

degree reduction in turning in the uncontrolled case. Results with 

B=1.0 (Fig 6b) again show little effect at the low frequencies and a 

significant reduction in losses and greater turning at the two higher 

frequencies. Increasing B further to 1.5 in Fig. 6c shows a further 

reduction in the loss peak magnitudes, and the effect is now visible 

even at F=0.28, particular with D=50%. With B increased to 2.0 and 

3.0 in Figs. 6d and 6e, there is not much change in the loss magnitude, 

but the flow turning increases by about 2 degrees in the best cases. 

 To summarize the results, the lift on the airfoil can be determined by 

integrating the difference between the suction and pressure side Cp

values along the axial direction. 

∫

− =

Cx

0 x

suction pressure

int C

Cp( dx)Cp Cp (1) 

This result is normalized with the result from the inviscid solution and 

shown as a function of B Fig. 7 for all the Re=25,000 cases. The VGJs 

are able to increase the lift by up to 20%, and as noted above, the 

effect increases with pulsing frequency and blowing ratio. If F=0.56 

or higher, B=0.75 is sufficient to reach the maximum lift attained. 

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0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

F=0.56 D=10

F=0.56 D=50

F=1.12 D=10

F=1.12 D=50

Inviscid

a)

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

F=0.14 D=50

F=0.28 D=2

F=0.28 D=10

F=0.28 D=60

F=0.56 D=10

F=0.56 D=50

F=1.12 D=10

F=1.12 D=50

Inviscid

b)

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

F=0.14 D=50

F=0.28 D=2

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

F=1.12 D=10

F=1.12 D=50

Inviscid

c)

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

Steady jets

F=0.14 D=2

F=0.14 D=10

F=0.14 D=50

F=0.28 D=2

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

F=1.12 D=10

F=1.12 D=50

Inviscid

d) 

Fig. 5 Cp profiles for Re=25,000 cases, B: a) 0.25, b) 0.5, c) 

0.75, d) 1.0, e) 1.5, f) 2.0, g) 2.5, h) 3.0 

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

F=0.14 D=50

F=0.28 D=2

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

F=1.12 D=10

F=1.12 D=50

Inviscid

e)

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

Steady jets

F=0.14 D=2

F=0.14 D=10

F=0.14 D=50

F=0.28 D=2

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

F=1.12 D=10

F=1.12 D=50

Inviscid

f)

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

Steady jets

F=0.14 D=50

F=0.28 D=2

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

F=1.12 D=10

F=1.12 D=50

Inviscid

g)

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/ls

cp

No jets

Steady jets

F=0.14 D=2

F=0.14 D=10

F=0.14 D=50

F=0.28 D=2

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

F=1.12 D=10

F=1.12 D=50

inviscid

h) 

Fig. 5 Cp profiles for Re=25,000 cases, B: a) 0.25, b) 0.5, c) 

0.75, d) 1.0, e) 1.5, f) 2.0, g) 2.5, h) 3.0 

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−1.5 −1 −0.5 0 0.5 1 1.5 0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4 B5 B6

No jets

F=1.12 D=10 B=0.25

F=0.56 D=10 B=0.5

F=1.12 D=50, B=0.5

F=0.56 D=10 B=0.75

a)

−1.5 −1 −0.5 0 0.5 1 1.5 0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4 B5 B6

No jets

Steady jets

F=0.14 D=50

F=0.28 D=10

F=0.56 D=10

F=1.12 D=10

b)

−1.5 −1 −0.5 0 0.5 1 1.5 0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4 B5 B6

No jets

F=0.28 D=2

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

c)

−1.5 −1 −0.5 0 0.5 1 1.5 0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4 B5 B6

No jets

Steady jets

F=0.14 D=50

F=0.28 D=10

F=0.56 D=10

F=1.12 D=10

d) 

Fig. 6 Total pressure loss profiles for Re=25,000 cases, B: 

a) 0.25-0.75, b) 1.0, c) 1.5, d) 2.0, e) 3.0 

−1.5 −1 −0.5 0 0.5 1 1.5 0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4 B5 B6

No jets

Steady jets

F=0.14 D=2

F=0.14 D=10

F=0.14 D=50

F=0.28 D=10

F=0.56 D=10

F=1.12 D=10

e) 

Fig. 6 Total pressure loss profiles for Re=25,000 cases, B: 

a) 0.25-0.75, b) 1.0, c) 1.5, d) 2.0, e) 3.0 

 To summarize the losses, the integrated loss around the center blade 

is computed as 

( )

( ) ST

Tej

1

j

2/L

2/L

int PP

PP

m

m

L

d

−

− = ∫ +

− &

φ &

φ φ

φψ ψ . (2) 

The first term on the right in Eq. (2) is an area averaged value as used 

in Bons et al. [25], as opposed to a mass average value as presented in 

Volino [6]. The mass and area average values are nearly equal for the 

present cases since the velocity deficit in the wake at 0.63Cx

downstream of the trailing edge is low. The second term on the right 

in Eq. (2) accounts for the pressure loss associated with the jet flow, as 

explained by Cully et al. [37]. The terms m1& , m j & and Pj are the main 

and jet mass flow rates and the cavity pressure. With steady jets, this 

jet associated loss term has a value of 0.21 when B=3, drops to about 

0.05 when B=2, and is below 0.01 for B=1 or lower. For pulsed jet 

cases, the jet loss term should be integrated over the pulsing cycle, but 

since the instantaneous cavity pressure is not measured, the term is 

estimated by multiplying the steady flow value by the duty cycle. This 

approximation overestimates the loss, since the jet velocity and mass 

flow rate drop after the initial peak, as shown in Figs. 2 and 4. Given 

the low jet mass flow rate, particularly at low B and D, the absolute 

value of the term is small and the approximation has a negligible effect 

on the overall ψint. The integrated loss is shown in Fig. 8. The VGJs 

clearly have a favorable effect, particularly at the higher frequencies. 

In the best case, ψint is reduced to about 40% of the baseline case level. 

With B at an achievable level of 1.0, ψint is reduced to about 60% of 

the baseline level. 

Re=50,000 

 Figure 9a shows Cp profiles for cases with Re=50,000 and B=0.25. 

The jets have no effect at F=0.14 (6 Hz), and some effect at F=0.28, 

which increases with duty cycle. With F=0.56 and D=50%, the 

separation bubble appears to close, although the plateau between 

s/Ls=0.55 and 0.8 indicates a large bubble is still present. When B is 

raised to 0.5, Fig. 9b again shows little effect with F=0.14, and a 

significant but closed bubble at the higher frequencies. As B is 

increased from 0.75 to 2.0 (Figs. 9c-9e), pulsing at F=0.14 begins to 

have an effect in reducing separation. With B=2.0, F=0.56, and 

D=50%, the separation bubble appears to be nearly eliminated, as in 

the B=2.0 steady blowing case. 

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0 0.5 1 1.5 2 2.5 3 0.8

0.85

0.9

0.95

1

1.05

1.1

B

Cpint/Cpinv

Steady jets

F=0.14, D=2

F=0.14, D=10

F=0.14, D=50

F=0.28, D=2

F=0.28, D=10

F=0.28, D=50

F=0.56, D=10

F=0.56, D=50

F=1.12, D=10

F=1.12, D=50

Fig. 7 Ratio of lift to lift in inviscid case, as indicated by 

integrated Cp result for Re=25,000 cases 

0 0.5 1 1.5 2 2.5 3 0

0.1

0.2

0.3

0.4

0.5

0.6

B

ψint

No jets, no holes

Steady jets

F=0.14 D=2

F=0.14 D=10

F=0.14 D=50

F=0.28 D=2

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

F=1.12 D=10

Fig. 8 Integrated total pressure loss for Re=25,000 cases 

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

Steady jets

F=0.14 D=10

F=0.14 D=50

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

Inviscid

a)

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

Steady jets

F=0.14 D=10

F=0.14 D=50

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

Inviscid

b) 

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

Steady jets

F=0.14 D=10

F=0.14 D=50

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

Inviscid

c)

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

Steady jets

F=0.14 D=10

F=0.14 D=50

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

Inviscid

d) 

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

Steady jets

F=0.14 D=10

F=0.14 D=50

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

Inviscid

e) 

Fig. 9 Cp profiles for Re=50,000 cases, B: a) 0.25, b) 0.5, c) 

0.75, d) 1.0, e) 2.0

 The corresponding loss coefficients are shown in Fig. 10. With 

B=0.25 in Fig. 10a, F=0.28 pulsing reduces losses below the steady 

case and pulsing with F=0.56 reduces losses further. The magnitude 

of the loss peaks at F=0.56 are about 80% of those in the best cases 

with steady blowing in Fig. 10e-f, and with much lower B. No affect 

of duty cycle is apparent in Fig. 10a. As B is increased from 0.5 to 2.0 

(Figs. 10b-10f), the loss results are consistent with the pressure 

profiles of Fig. 9. With F=0.56, the losses are already low with 

B=0.25, so increasing B further has a relatively small effect. With 

F=0.14 and 0.28, the losses drop as B and D are increased, eventually 

approaching the F=0.56 results. In all cases pulsed blowing reduces 

losses compared to steady blowing. Even in the best case, however, 

the loss peaks are still about 30% higher than in the high Re≥200,000 

baseline cases [6]. 

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−1.5 −1 −0.5 0 0.5 1 1.5 0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4 B5 B6

No jets

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

a)

−1.5 −1 −0.5 0 0.5 1 1.5 0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4 B5 B6

No jets

F=0.14 D=50

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

b)

−1.5 −1 −0.5 0 0.5 1 1.5 0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4 B5 B6

No jets

Steady jets

F=0.14 D=50

F=0.28 D=10

F=0.28 D=50

c)

−1.5 −1 −0.5 0 0.5 1 1.5 0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4 B5 B6

No jets

Steady jets

F=0.14 D=10

F=0.14 D=50

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

d)

−1.5 −1 −0.5 0 0.5 1 1.5 0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4 B5 B6

No jets

steady jets

F=0.14 D=10

F=0.14 D=50

F=0.28 D=10

F=0.28 D=50

e)

−1.5 −1 −0.5 0 0.5 1 1.5 0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4 B5 B6

No jets

Steady jets

F=0.14 D=10

F=0.28 D=10

F=0.28 D=50

f) 

Fig. 10 Total pressure loss profiles for Re=50,000 cases, B: 

a) 0.25, b) 0.5, c) 0.75, d) 1.0, e) 1.5, f) 2.0 

 The Cp results for the Re=50,000 cases are summarized in Fig. 11. 

As at Re=25,000, the VGJs are able to increase the lift by about 20% 

over the baseline case, and are more effective at higher B and F. The 

jets are most effective for F≥0.28, even with B=0.25 and D=10%. The 

integrated total pressure loss in Fig. 12 shows that losses can be 

reduced by about 55% below baseline levels. With F=0.28, B=1 to 1.5 

is needed to achieve this reduction, but with F=0.56, the same 

reduction is possible with B=0.25. 

Re=100,000 

 In the baseline case at Re=100,000, the boundary layer separated 

and did not reattach. With the VGJ holes present, even without 

blowing, the shear layer was on the verge of reattachment. Pressure 

coefficients are shown in Fig. 13. All cases show boundary layer 

reattachment, even with B=0.25, F=0.14 and D=10%. At this 

Reynolds number, almost any small disturbance is sufficient to cause 

the boundary layer to reattach. Once it is attached, it tends to stay 

attached even if the flow control is turned off, and will only 

occasionally separate again. Raising B and the pulsing frequency does 

help to reduce the size of the separation bubble slightly, as indicated 

by the pressure coefficients. 

 Loss profiles for some of the Re=100,000 cases are shown in Fig. 

14. It appears that with B=0.5 or greater for steady blowing or B=025 

or greater with pulsed jets, the total pressure loss is driven to its lowest 

possible value at this Reynolds number. This is well below the 

uncontrolled value at this Re, but the peaks are about 30% higher than 

in baseline cases with Re≥200,000 [6]. Even with the best VGJ flow 

control at Re=100,000, it appears a small separation bubble is still 

present, which increases the boundary layer thickness and losses 

compared to higher Reynolds number cases. 

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0 0.5 1 1.5 2 0.8

0.85

0.9

0.95

1

1.05

1.1

B

Cpint/Cpinv

Steady jets

F=0.14 D=10

F=0.14 D=50

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

Fig. 11 Ratio of lift to lift in inviscid case, as indicated by 

integrated Cp result for Re=50,000 cases

0 0.5 1 1.5 2 2.5 3 0

0.1

0.2

0.3

0.4

0.5

0.6

B

ψint

No jets, no holes

Steady jets

F=0.14 D=10

F=0.14 D=50

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

Fig. 12 Integrated total pressure loss for Re=50,000 cases

 The integrated Cp and loss are shown in Figs. 15 and 16. As in the 

lower Re cases, the lift rises about 20% above the baseline value in all 

cases with flow control. The total pressure loss drops by about 70%. 

NUMERICAL RESULTS 

 Although the Transition-sst model predicted the baseline cases well, 

its performance is not good in cases with VGJ flow control. The 

model fails to predict reattachment, so the predictions of the pressure 

coefficients are still generally satisfactory when the experiments do 

not show reattachment. Some examples of Cp results are shown in 

Fig. 17 for cases at Re=50,000. In the baseline case and with steady 

jets at B=0.5, both the experiments and the CFD show separation 

without reattachment, with little change due to the VGJs. Raising B to 

2.0 with steady blowing, the experiment shows clear reattachment. In 

fact, steady blowing with B=1.0 was shown to be sufficient for 

reattachment in Fig. 9. The CFD correctly shows a rise in the Cp peak 

and some drop toward the trailing edge, but the effect of the VGJs is 

much weaker than in the experiment, and reattachment is not 

predicted. With pulsed jets at B=0.5, F=0.28 and D=10%, the 

experiment shows reattachment after a large separation bubble. The 

CFD shows no difference from the steady blowing case at B=0.5 and 

does not predict reattachment. Results are similar at the other 

Reynolds numbers. At Re=25,000 with B=1, F=0.56 and D=10%, the 

experiment shows reattachment after a large separation bubble, but the 

CFD indicates no reattachment and essentially no change in Cp from 

the baseline case. At Re=100,000, cases were tried with steady 

blowing at B=0.25 and 1.0 and with pulsed jets at B=0.25, F=0.14 and 

D=10%. All of these cases show a slight rise in the Cp maximum. 

The CFD cases all agree with each other, showing no additional effect 

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

Steady jets

F=0.14 D=10

F=0.14 D=50

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

Inviscid

a)

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

Steady jets

F=0.14 D=10

F=0.14 D=50

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

Inviscid

b)

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

Steady jets

F=0.14 D=10

F=0.14 D=50

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

Inviscid

c)

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

Steady jets

F=0.14 D=10

F=0.14 D=50

F=0.28 D=10

F=0.28 D=50

F=0.56 D=50

Inviscid

d) 

Fig. 13 Cp profiles for Re=100,000 cases, B: a) 0.25, b) 0.5, 

c) 0.75, d) 1.0

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distribution is unlimited. 

−1.5 −1 −0.5 0 0.5 1 1.5 0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4 B5 B6

No jets, no holes

No jets, w/holes

B=0.25, steady

B=0.50, steady

B=0.75, steady

B=1.0, steady

F=0.14 D=50 B=0.25

F=0.14 D=50 B=0.75

Fig. 14 Total pressure loss profiles for Re=100,000 cases 

0 0.2 0.4 0.6 0.8 1 0.8

0.85

0.9

0.95

1

1.05

1.1

B

Cpint/Cpinv

Steady jets

F=0.14 D=10

F=0.14 D=50

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

Fig. 15 Ratio of lift to lift in inviscid case, as indicated by 

integrated Cp result for Re=100,000 cases 

0 0.2 0.4 0.6 0.8 1 0

0.1

0.2

0.3

0.4

0.5

B

ψint

No jets, no holes

Steady jets

F=0.14 D=50

Fig. 16 Integrated total pressure loss for Re=100,000 cases

of the pulsing or increased B. As in the Re=50,000 cases, the effect of 

the VGJs is much smaller than in the experiments, and the 

reattachment is missed. 

 Although the Transition-sst model produced good results in the 

baseline cases, it was clearly unable to capture the effect of the VGJs. 

The VGJs must produce flows with are locally three dimensional and 

the resulting local turbulence may be highly anisotropic. Perhaps a 

higher order model may be able to capture these effects. The moderate 

success that others such as Rizzeta and Visbal [34] had with LES 

suggest that additional grid refinement away from the wall, where the 

jets are interacting with the main flow may also help. 

0 0.2 0.4 0.6 0.8 1

0

0.5

1

1.5

2

2.5

3

s/Ls

Cp

No jets E

No jets C

B=0.50 steady E

B=0.50 steady C

B=2.0 steady E

B=2.0 steady C

B=0.5 F=0.56 D=10 E

B=0.5 F=0.56 D=10 C

Inviscid

Fig. 17 Comparison of experimental (E) and CFD (C) Cp 

profiles for Re=50,000 cases

DISCUSSION 

 When Re=100,000, the boundary layer is on the verge of 

reattachment even without the jets, so any jet blowing causes 

reattachment, an increase in lift and a large reduction in losses, 

independent of blowing ratio, frequency or duty cycle. At the lower 

Reynolds numbers, the effect of varying the VGJ parameters is more 

significant. It appears that pulsing with F=0.14 is not effective unless 

the blowing ratio is so strong and the duty cycle so long that the results 

begin to approach the steady blowing case. Pulsing with F=0.28 is 

more effective and is duty cycle dependent. It appears that 

intermittent control may be achieved at this frequency with periods of 

no control between pulses. Increasing the duty cycle may help to 

lengthen the period of separation control, as does increasing the 

blowing rate. Pulsing with F=0.56 or higher results in flow control 

even with a low duty cycle and moderate blowing ratio of B=0.75 in 

the Re=25,000 case, and with B=0.25 in the Re=50,000 case. The 

favorable results at low duty cycle indicate that the initial pulse is able 

to induce boundary layer reattachment, and the effect is sustained long 

enough when the jets are off to maintain some control until the next 

pulse. 

 Volino [15] found that the trailing edge of the calmed region 

following a VGJ disturbance travels at about 0.3 times the freestream 

velocity. This suggests that if the VGJs are pulsed at F=0.3 or higher, 

a disturbance or its calmed region would always be present in the 

boundary layer. At lower frequencies there would be periods between 

pulses when the boundary layer could relax to its uncontrolled state 

and separate. The present results agree with this, showing more 

effective control when F is above 0.28 than when it is below. If the 

boundary layer does begin to relax between pulses, the amount of 

relaxation necessary before separation occurs would likely depend on 

the pressure gradient and Reynolds number. This may explain why 

pulsing at F≤0.28 is more effective as Re is increased from 25,000 to 

100,000, and why Bons et al. [18] were able to able to suppress 

separation with F=0.1 and very low duty cycle on the less aggressive 

Pack B airfoil. 

CONCLUSIONS 

 The effect of vortex generator jets on the flow over the very high lift 

L1A airfoil was studied experimentally and computationally under low 

freestream turbulence conditions. Reynolds numbers based on suction 

surface length and nominal exit velocity of 25,000, 50,000 and 

100,000 were considered. Without flow control, the boundary layer 

separated in all cases and did not reattach. Flow control with VGJs 

NASA/CR—2012-217415 74
12 

This material is declared a work of the U.S. Government and is not subject to copyright protection in the United States. Approved for public release; 

distribution is unlimited. 

was possible even at the lowest Reynolds numbers. In agreement with 

previous studies, pulsed jets were found more effective than steady 

jets. A pulsing frequency of F=0.28 was marginal for good control at 

moderate blowing ratios. Pulsing with F=0.56 or higher allowed for 

separation control even with blowing ratios as low as 0.25 and duty 

cycles of 10%. Effective separation control resulted in a 20% increase 

in lift and up to a 70% reduction in total pressure loss compared to 

baseline cases at the same Reynolds number. 

 The Transition-sst model, which showed promising results in the 

cases without flow control, did not predict reattachment in the URANS 

calculations in any cases with VGJs. A finer grid away from the wall 

or a higher order model which could better capture the three 

dimensionality and anisotropy of the jets might give better results. 

 The present experimental results show the parameter ranges where 

VGJs can be effective with the L1A airfoil. Further measurements of 

instantaneous velocity, both in the boundary layer and wake may be 

helpful for clarifying the physics of how the VGJs work. Such 

knowledge may be useful for developing improved computational 

models. 

ACKNOWLEDGMENTS

 This work was sponsored by the National Aeronautics and Space 

Administration. The grant monitor is Dr. Anthony Strazisar of the 

NASA Glenn Research Center. The support of the United States 

Naval Academy Technical Support Department Shop and Fluids 

Laboratory is greatly appreciated. 

 

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Loaded Low-Pressure Turbine Airfoil,” ASME Paper GT2008-

51445.

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Low-Pressure Turbine Airfoil Conditions: Part 1 – Mean Flow 

and Turbulence Statistics,” ASME Journal of Turbomachinery, 

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13. 

 

[11] Zhang, X.F., Vera, M., Hodson, H., and Harvey, N., 2007, 

“Separation and Transition Control on an Aft-Loaded Ultra-HighLift LP Turbine Blade at Low Reynolds Numbers: Low-Speed 

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Sondergaard, R., 2005, “Designing Low-Pressure Turbine Blades 

With Integrated Flow Control,” ASME Paper GT2005-68962 

[13] Bohl, D.G., and Volino, R.J., 2006, “Experiments with ThreeDimensional Passive Flow Control Devices on Low-Pressure 

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251-260. 

[14] Volino, R.J., 2003, “Passive Flow Control on Low-Pressure 

Turbine Airfoils,” ASME Journal of Turbomachinery, 125, pp. 

754-764. 

[15] Volino, R.J., 2003, “Separation Control on Low-Pressure Turbine 

Airfoils Using Synthetic Vortex Generator Jets,” ASME Journal 

of Turbomachinery, 125, pp. 765-777. 

[16] Huang, J., Corke, T., and Thomas, F., 2003, “Plasma Actuators 

for Separation Control on Low Pressure Turbine Blades,” AIAA 

Paper 2003-1027. 

[17] Johnston, J.P., and Nishi, M., 1990, “Vortex Generator Jets. 

Means for Flow Separation Control,” AIAA Journal, 28, pp. 989-

994. 

[18] Bons, J.P., Sondergaard, R., and Rivir, R.B., 2002, “The Fluid 

Dynamics of LPT Blade Separation Control Using Pulsed Jets,” 

ASME Journal of Turbomachinery, 124, pp. 77-85. 

[19] Volino, R.J., 2003, “Separation Control on Low-Pressure Turbine 

Airfoils Using Synthetic Vortex Generator Jets,” ASME Journal 

of Turbomachinery, 125, pp. 765-777. 

[20] Volino, R.J., and Bohl, D.G., 2005, “Structure of Oscillating 

Vortex Generator Jets,” Proceedings of the Fourth International 

Symposium on Turbulence and Shear Flow Phenomena, 2, pp. 

589-594.

[21] McQuilling, M., and Jacob, J., 2004, “Effect of Chord Location 

on Separation Control With Vortex Generator Jets on Low 

Pressure Turbine Blades,” AIAA Paper 2004-2205. 

[22] Eldredge, R. G., and Bons, J. P., 2004, “Active Control of a 

Separating Boundary Layer With Steady Vortex Generating 

Jets—Detailed Flow Measurements,” AIAA Paper 2004-751. 

[23] Gostelow, J.P., Walker, G.J., Solomon, W.J., Hong, G., and 

Melwani, N., 1997, “Investigation of the Calmed Region Behind 

a Turbulent Spot,” ASME Journal of Turbomachinery, 119, pp. 

802-809. 

[24] Schulte, V., and Hodson, H.P., 1998, “Prediction of the Becalmed 

Region for LP Turbine Profile Design,” ASME Journal of 

Turbomachinery, 120, pp. 839-846. 

[25] Bons, J.P., Reimann, D., and Bloxham, M., 2008, “Separated 

Flow Transition on an LP Turbine Blade With Pulsed Flow 

Control,” ASME Journal of Turbomachinery, 130, 021014. 

[26] Clark J.P., 2007, Private Communication, Air Force Research 

Laboratory. 

[27] Zhang, X.F., and Hodson, H., 2005, “Combined Effects of 

Surface Trips and Unsteady Wakes on the Boundary Layer 

Development of an Ultra-High-Lift LP Turbine Blade,” ASME 

Journal of Turbomachinery, 127, pp. 479-488. 

[28] Ibrahim, M., Kartuzova, O., and Volino, R.J., 2008, 

“Experimental and Computational Investigations of Separation 

and Transition on a Highly Loaded Low Pressure Turbine Airfoil: 

Part 1 – Low Freestream Turbulence Intensity,” ASME Paper 

IMECE2008-68879. 

NASA/CR—2012-217415 75
13 

This material is declared a work of the U.S. Government and is not subject to copyright protection in the United States. Approved for public release; 

distribution is unlimited. 

 

[29] Volino, R.J, Kartuzova, O., and Ibrahim, M., 2008, 

“Experimental and Computational Investigations of Separation 

and Transition on a Highly Loaded Low Pressure Turbine Airfoil: 

Part 2 – High Freestream Turbulence Intensity,” ASME Paper 

IMECE2008-68776.

[30] Suzen, Y.B., Huang, P.G., Ashpis, D.E., Volino, R.J., Corke, 

T.C., Thomas, F.O., Huang, J., Lake, J.P. and King, P.I., 2007, 

“A Computational Fluid Dynamics Study of Transitional Flows 

in Low-Pressure Turbines Under a Wide Range of Operating 

Conditions,” ASME Journal of Turbomachinery, 129, pp. 527-

541. 

[31] Menter, F.R., Langtry, R.B., Likki, S.R., Suzen, Y.B., Huang, 

P.G., and Völker, S., 2006, “A Correlation Based Transition 

Model Using Local Variables – Part I: Model Formulation”, 

ASME Journal of Turbomachinery, 128, pp. 413-422. 

[32] Langtry, R.B., Menter, F.R., Likki, S.R., Suzen, Y.B., Huang, 

P.G., and Völker, S., 2006, “A Correlation based Transition 

Model using Local Variables Part II: Test Cases and Industrial 

Applications,” ASME Journal of Turbomachinery, 128, pp. 423-

444. 

[33] Gross, A., Fasel, H.F., 2008 “Strategies for Simulating Flow 

Through Low-Pressure Turbine Cascade,” ASME Journal of 

Fluids Engineering, 130, 111105. 

[34] Rizzetta, D.P., and Visbal, M.R., 2005, “Numerical Simulation of 

Separation Control for Transitional Highly Loaded Low-Pressure 

Turbines,” AIAA Journal, 43, pp.1958-1967. 

[35] Garg V.K., 2002 “Low-Pressure Turbine Separation Control - 

Comparison with Experimental Data,” ASME Paper GT-2002-

30229. 

[36] Menter, F.R., 1994, “Two-Equation Eddy-Viscosity Turbulence 

Models for Engineering Applications,” AIAA Journal, 32, pp. 

1598-1605. 

[37] Culley, D.E., Bright, M.M., Prahst, P.S., and Stazisar, A.J., 2004, 

“Active Flow Separation Control of a Stator Vane Using 

Embedded Injection in a Multistage Compressor Experiment,” 

ASME Journal of Turbomachinery, 126, pp. 24-34. 

NASA/CR—2012-217415 76
1 

Final draft. 

Submitted for publication in: K. Hanjalić, Y. Nagano and S. Jakirlić (Editors), “Turbulence, Heat and 

Mass Transfer 6,” Begell House, Inc., 2009. 

 

LES and URANS Computational Investigations of 

LPT Blade (L1A) Separation Control using

Vortex Generator Jets

M. B. Ibrahim1, O. Kartuzova2 and R. J. Volino3

1

Department of Mechanical Engineering, Cleveland State University, 

2121 Euclid Ave., SH232, Cleveland, OH 44115, m.ibrahim@csuohio.edu

2

Department of Mechanical Engineering, Cleveland State University, 

2121 Euclid Ave., SH232, Cleveland, OH 44115, kartuzova_olga@hotmail.com

3

Mechanical Engineering Department, United States Naval Academy, 

Annapolis, Maryland 21402-5042, volino@usna.edu

Abstract - This paper describes active flow control utilizing steady VGJs both 

experimentally and computationally for the L1A airfoil. URANS approach is compared 

with LES to test the ability to accurately predict effect of VGJs on the boundary layer 

separation. Cases were considered at Reynolds numbers (based on the suction surface 

length and the nominal exit velocity from the cascade) of 25,000, 50,000 and 100,000. In 

all cases without flow control, the boundary layer separated and did not reattach. The VGJs 

were successful in reducing the bubble size or removing it completely depending on the Re 

number and blowing ratio (B). Transition-sst RANS model was successful in predicting the 

flow separation and attachments with no jet flow. LES showed superior performance for the 

VGJs cases and much better agreement with the experimental data for the flow field at 

different values of Re and B.

1. Nomenclature

B blowing ratio, maximum jet velocity/local freestream velocity

Cp ( ) 2

T − ρU/PP2 e , pressure coefficient

Cx axial chord length

D jet diameter (=0.8 mm)

Ls suction surface length

Lφ blade spacing (pitch)

P pressure

PT upstream stagnation pressure 

Re UeLs/ν, exit Reynolds number 

Reθt transition momentum thickness Reynolds number 

s streamwise coordinate, distance from leading edge

u' RMS of the fluctuating component of the streamwise velocity

U streamwise velocity

Ue nominal exit freestream velocity, based on inviscid solution 

NASA/CR—2012-217415 77
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Vx velocity in the axial direction

x axial distance from leading edge

y distance from the wall

z distance in the spanwise direction

ν kinematic viscosity

ρ density 

2. Introduction 

Boundary layer separation on the suction side of low-pressure turbine (LPT) airfoils 

can occur due to strong adverse pressure gradients. The problem is becoming more severe as 

airfoil loading is increased. If the boundary layer separates, the lift from the airfoil decreases 

and the aerodynamic loss increases, resulting in a drop in overall engine efficiency. A 

significant increase in efficiency could be achieved if separation could be prevented, or 

minimized. Active flow control could provide a means for minimizing separation under 

conditions where it is most severe (low Re), without causing additional losses under other 

conditions (high Re). Minimizing separation will allow improved designs with fewer stages 

and fewer airfoils per stage to generate the same power. The active flow control technique 

called Vortex Generator Jets (VGJs), introduced by Johnston and Nishi [1], is the subject of 

this paper. In this technique blowing from small, compound angled holes is used to create 

streamwise vortices. The vortices bring high momentum fluid into the near wall region as 

well as promote transition and turbulent mixing. This helps to control separation. A review of 

experimental work on this topic was given in Volino et al. [2]. 

Along with the experimental investigations, numerical simulations of the flow over 

LPT blades, utilizing steady and pulsed vortex generator jets (VGJs) were performed by 

different researchers. This type of flow is challenging for CFD because of its transitional 

nature in combination with highly 3D jet flow. 

Rizzetta and Visbal [3] used ILES (Implicit Large Eddy Simulation) to investigate the 

effect of flow control by pulsed VGJs on the flow separation in a Pack B cascade. The Pack 

B is a widely studied LPT airfoil. They reported that for inlet Re = 25,000 and B=2 flow 

control helped to keep the flow attached for an additional 15% of the chord. Despite some 

differences with experiment, numerical and experimental time-mean velocity profiles were in 

reasonable agreement. 

Postl et al. [4] studied the effect of active flow control utilizing steady and pulsed 

VGJs on preventing laminar separation on the Pack B airfoil. In their study they used two 

computational approaches. The first was direct numerical simulation (DNS) of the flow over a 

flat plate with imposed streamwise pressure gradient as measured on the suction side of the 

Pack B airfoil. In the second approach a linear LPT cascade was simulated with the flow 

control cases performed in 2D using slots. They observed two different physical mechanisms 

to affect the control for the steady versus pulsed VGJs. The steady jets developed streamwise 

structures which enhanced momentum exchange between the near wall region and free 

stream. For the pulsed VGJs accelerated boundary layer transition helped to achieve flow 

reattachment. Their results are in reasonably good agreement with experimental data except

in the separated region, where the size of the separation bubble was under predicted. 

Garg [5] used the NASA Glenn-HT code with the k-ω SST model of Menter [6] to 

compute the flow over the Pack B blade with and without use of VGJs. This work resulted in 

correct predictions of the separation location in the baseline case (without VGJs) as well as 

showing that separation vanishes in the flow control case as in experiment. However, the 

separated region and the wake were not well predicted, which is common for RANS.

In the present study the very highly loaded L1A airfoil was considered. The L1A was 

designed at the Air Force Research Laboratory (AFRL) and is available on a limited basis 

NASA/CR—2012-217415 78
3 

from Clark [7]. It is an aft loaded blade with the same flow angles and loading as the mid 

chord loaded L1M. Based on the design calculations of Clark, [7], the L1A has 10% higher 

loading than the “ultra-high lift” airfoils described by Zhang and Hodson [8], and 17% higher 

loading than the Pack B. Because the L1A is aft loaded, it is more prone to separation than the 

L1M, as documented in Bons et al. [9], Ibrahim et al. [10], and Volino et al. [11]. In cases 

without flow control and with low freestream turbulence, the boundary layer separates when 

Re<150,000 and does not reattach, in spite of transition to turbulence in the shear layer over 

the separation bubble in all cases. This result contrasts with the results of earlier studies on 

less aggressive airfoils, which all showed reattachment after transition. The separation bubble 

on the L1A is about four times thicker than that on the Pack B. The larger distance from the 

shear layer to the wall on the L1A apparently prevents the turbulent mixing in the shear layer 

from reaching the wall and causing reattachment. The failure of the boundary layer to reattach 

results in a 20% loss in lift and increases profile losses by a factor of 7 [2]. At higher 

Reynolds numbers the separation bubble is small and the boundary layer is attached over 

most of the airfoil. In cases with high freestream turbulence, results are similar, but the shear 

layer is somewhat thicker and the separation bubble thinner due to increased mixing induced 

in the shear layer. This results in reattachment after transition at Re=50,000 and 100,000. At 

the lowest Re considered (25,000) the boundary layer still does not reattach. 

Boundary layer separation control has been studied using Vortex Generator Jets 

(VGJs) on the L1A airfoil. Experiments were done under low freestream turbulence 

conditions on a linear cascade in a low speed wind tunnel (see Volino et al., [2]). In that paper 

the pressure surveys on the airfoil surface and downstream total pressure loss surveys were 

documented. Cases were considered at Reynolds numbers (based on the suction surface 

length and the nominal exit velocity from the cascade) of 25,000, 50,000 and 100,000. In all 

cases without flow control, the boundary layer separated and did not reattach. These cases 

(without jets) were documented experimentally and computationally for low freestream 

turbulence intensity (see Ibrahim et al. [10]) and high freestream turbulence intensity (see 

Volino et al. [11]).

This paper describes active flow control utilizing steady VGJs both experimentally 

and computationally for the L1A airfoil. A previously developed computational model for the 

baseline cases (without VGJs), by the authors, [10,11] is enhanced by additional grid 

independence study for the necessary grid resolution around the jets and in the spanwise 

direction. Unsteady RANS approach is compared with LES to test the ability to accurately 

predict the effect of VGJs on the boundary layer separation. 

3. Numerical Simulation

The numerical simulations were conducted for the L1A airfoil utilizing version 6.3.26 

of the finite-volume code FluentTM[12]. Cases (with jet) were considered, in this paper, at

different Reynolds numbers and blowing ratios. Table 1. shows a summary of all the cases 

presented in this paper both experimentally and computationally. Comparison for the results 

of these cases are shown in Section 5. Due to high flow unsteadiness even for steady blowing, 

unsteady calculations were done for all cases. The three dimensional computational domain 

includes a single passage (see Table 2. for cascade parameters). A uniform velocity inflow 

condition is specified 1.9 Cx upstream of the blade leading edge in the flow direction. The 

inlet flow angle is set to 33o based on an inviscid calculation of the full cascade used in the 

experiment [10]. This angle agrees with the experimentally measured inlet angle to within the 

experiment uncertainty. The exit boundary is located 3.8 Cx downstream of the trailing edges 

in the flow direction. In the spanwise direction, the domain includes one VGJ. The boundary 

conditions on the sides of the passage are periodic. The full length of the hole is included in 

the simulations, allowing the jet velocity profile to develop before entering the main domain 

[2]. A uniform velocity boundary condition is specified at the hole inlet. 

NASA/CR—2012-217415 79
4 

Convergence was established when: 1) residuals reduced to a value 10-5, 2) no change was 

observed in any field results, and 3) the mass imbalance was less than 0.01 %. 

Table 1. Test matrix used in experiments and CFD (NA = not available). 

Reynolds number, Re 25,000 50,000 100,000

Blowing ratio, B 1 3 0.5 2 0.25 0.75 1

Cp

Experiment      NA 

LES      NA 

Transition-sst NA  NA  NA NA 

U/Ue

Experiment NA  NA  NA  NA

LES      NA 

Transition-sst NA  NA  NA NA 

u'/Ue

Experiment NA  NA  NA  NA

LES NA  NA  NA NA 

Transition-sst NA

Table 2. Cascade parameters. 

Axial Chord, 

Cx [mm]

True Chord 

[mm]

Pitch, Lφ

[mm]

Span

[mm]

Suction side, 

Ls [mm]

Inlet flow 

angle

Exit flow 

angle

134 146 136 724 203 35° 60°

4. Turbulence Models Description

4.1. Transition-sst (4 eq.) model

A new correlation-based transition model was proposed by Menter et al. [13]. This 

model is based on two transport equations. The intermittency transport equation is used to 

trigger the transition onset. The transport equation for the transition momentum thickness 

Reynolds number (Reθt) is used to capture non-local effects of freestream turbulence intensity 

and pressure gradient at the boundary layer edge. Outside the boundary layer the transport 

variable was forced to follow the value of Reθt given by correlations. Those two equations 

were coupled with the shear stress transport turbulence model (SST). This model has recently 

become available in Fluent code [12].

4.2. Large Eddy Simulation with Dynamic Kinetic Energy Subgrid-Scale model

In LES, large eddies are resolved directly, while small eddies are modeled. The 

rationale behind LES are: a) Momentum, mass, energy, and other passive scalars are 

transported mostly by large eddies, b) Large eddies are more problem-dependent; they are 

dictated by the geometries and boundary conditions of the flow involved, c) Small eddies are 

less dependent on the geometry, tend to be more isotropic, and are consequently more 

NASA/CR—2012-217415 80
5 

universal, and d) The chance of finding a universal turbulence model is much higher for small 

eddies.

The governing equations employed for LES are obtained by filtering the timedependent Navier-Stokes equations in either Fourier (wave-number) space or configuration 

(physical) space. The filtering process effectively filters out the eddies whose scales are 

smaller than the filter width or grid spacing used in the computations. The resulting equations 

thus govern the dynamics of large eddies. The subgrid-scale stresses resulting from the 

filtering operation are unknown, and require modeling. The subgrid-scale turbulence models 

in Fluent employ the Boussinesq hypothesis as in the RANS models. 

The dynamic subgrid-scale kinetic energy model in Fluent is based on the model 

proposed by Kim and Menon [14]. In this model a separate transport equation is solved for 

subgrid-scale kinetic energy. The model constants are determined dynamically. The details of 

the implementation of this model in Fluent and it's validation are given by Kim [15]. 

5. Results and Discussion

5.1. Code Validation 

 Three different grids were designed for this study as shown in Table 3. Figure 1shows 

the grid structure in the vicinity of the jet for Grid#37. 

 Table 3. Grid#3, 35 and 37 used in this investigation. 

Grid # Size (Cells) Number of grids in z 

direction

y+ ∆z+ ∆x+

3 1,500,000 15 0.5 12.6 1 – 100

35

(airf-mid)

5,900,000 30 0.5 6.3 0.4 – 52

37

(airf-mid)

11,900,000 54 0.5 0.4 - 3.5 0.4 – 52

To accurately represent structures in the near-wall region (for LES) recommended 

values are: y+~ 2; ∆x+ ~ 50-150; ∆z+ ~ 15-40 (see Piomelli and Chasnov [16]). Based on 

results (not shown for space limitation) for the pressure coefficient versus the dimensionless 

location on the suction side s/Ls, Grid#37 showed closet agreement to the data and therefore 

was chosen for further computation. This Grid#37 was then run for different time steps 

(0.0005, 0.0001 and 0.00005 s) and time step =0.0001 s was selected since no much 

improvement were achieved using the smaller one (0.00005 s). It should be noted that grid 

independence tests were conducted for the Transition-sst model without jet flow (see Ibrahim 

et al. [10]). In this study further grid refinement was conducted for two reasons: a) to handle 

highly 3D jet flow and b) to utilize LES computation. 

Velocity measurements were conducted at 6 different stations downstream of the 

suction peak (see Table 4.). Comparison will be made between the velocity profiles at these 

stations from CFD and experiments. 

Table 4. Velocity profile measurement stations. 

Station 1 2 3 4 5 6

s/Ls 0.53 0.59 0.69 0.78 0.88 0.97

x/Cx 0.65 0.72 0.80 0.86 0.92 0.97

NASA/CR—2012-217415 81
6 

5.2. Re - 25,000 steady VGJs cases 

 Pressure coefficient plotted versus dimensionless distance along the suction side of an 

airfoil is presented in Fig. 2. The experiment for B=1 shows separation at s/Ls between 0.53 

and 0.59, with no reattachment. At B=2 there is no clear indication if the flow is attached

along the airfoil since there seems to be a bubble present. Computational (LES) results for 

two blowing ratios (B = 1 and 3) are compared to the experiment with the same blowing 

conditions. For B = 1 both LES and experiment show flow separation starting after the 

suction peak with no reattachment, which is indicated by the large plateau in Cp after the 

suction peak. Cp at the suction peak is lower in CFD compared to experiment, but it is within 

the range of experimental uncertainty (0.07). The CFD results from the Transition-sst model 

(B=3) are shown in Fig. 2 also for comparison between RANS and LES. The Transition-sst 

model predicts Cp similar to LES, except for downstream locations (s/Ls = 0.8 - 1.0), where it 

predicts lower Cp values, than LES. This indicates a smaller separation bubble modeled by 

RANS than by LES for Re = 25,000 and B = 3. The reason for not showing results from

Transition-sst model and B = 1 is that at lower blowing ratio jets have no effect on separation 

and conditions are close to the "no-jets" cases. We know from our previous work [10,11] that 

the Transition-sst model predicts Cp reasonably well compared to experiment for "no-jets" 

cases.

 Velocity profiles normalized by the nominal exit velocity are plotted versus 

dimensionless distance from the wall in the direction normal to the wall in Fig. 3. The results 

for the 6 measurement stations (see Table 4.) located downstream of the suction peak of the 

airfoil are shown. Computational velocity profiles from LES for B = 1 and B = 3 are shown. 

Experimental data and Transition-sst results are only available for B = 3. For B = 1, LES 

shows separation starting at station 2 and large separation bubble is present at all stations 

from 2 to 6, based on negative velocities near the wall at those locations. For B = 3 both 

Transition-sst and LES show separation started at station 4 and continuing at stations 5 and 6. 

Separation bubble is smaller than that for B = 1. LES is in reasonably good agreement with 

experiment. It should be noted the larger disagreement near the wall is due to the fact that hot 

wire anemometry was used in the experiment, which is not capable of measuring negative 

velocities. The Transition-sst model underpredicts velocities near the wall and the size of the 

bubble, compared to LES and experiment. 

 Profiles of the streamwise component of the RMS fluctuating velocity, u', normalized 

by the nominal exit velocity are plotted versus dimensionless distance from the wall in Fig. 4. 

The results for 6 measurement stations located downstream of the suction peak of the airfoil 

are shown. Computational and experimental profiles are for B = 3. The experiment shows the 

location of the peak is away from the wall indicating the presence of a bubble. LES compares 

reasonably well with data including magnitude and location of u'.

For the purpose of visualization of the separated region as well as the influence of the 

jet's blowing ratio, iso-surfaces of instantaneous axial velocity Vx = 0.01 m/s are shown in 

Fig. 5 for B = 1 and B = 3. These CFD results are from the LES model. The reason for 

choosing Vx = 0.01 m/s is that this small (but not negative) value represents velocity in the 

shear layer of the separation bubble and helps to visualize the size of the bubble and shapes of 

the vortices created by the jets. The airfoil with 3 jets on the suction side near the suction 

peak is shown. In the case with B = 1 the low velocities in the shear layer of the separation 

bubble are at a distance from the wall, thus the separation bubble is large. It looks like not 

enough mixing is happening at this iso-surface and the fluid issuing from the jet is moving 

aligned with the cross flow direction at this low blowing ratio. The separation bubble is 

smaller in the B = 3 case, low velocities in the shear layer are closer to the wall than in the B 

= 1 case. The visualization shows more mixing happening near the wall (the iso-surface is 

NASA/CR—2012-217415 82
7 

less smooth than in B = 1 case). The fluid coming from the jet is moving at an angle with the 

flow direction downstream of the suction peak of the airfoil due to the high momentum jets at 

B = 3.

 Instantaneous axial vorticity contours are presented in Fig. 6 for the B = 1 and B = 3 

cases for the LES model. For B = 1 maximum vorticity is in the shear layer of the separation 

bubble away from the wall between stations 3 and 4. In the B = 3 case, the region with high 

vorticity (more mixing) is in the shear layer, but closer to the wall than in the B = 1 case. Its 

location is moved upstream and is between stations 2 and 3. More mixing in the high blowing 

ratio case (indicated by the streamwise vorticity contours) helps to reduce the size of the 

separation bubble (indicated also by the velocity plots - Fig. 3).

Fig. 7 presents a plane view of the 6 measurement stations (see Table 4.) along the 

airfoil suction side downstream of the jets locations. Subgrid turbulent kinetic energy (TKE) 

contours are displayed for B = 1 (7a) and B = 3 (7b). 

In summary, for B = 1: a) separation starts between stations 1 and 2, b) transition to 

turbulence starts between stations 3 and 4 (max subgird TKE and X-vorticity locations). For 

B = 3: a) separation starts near station 3 (see Figs. 2, 3 and 4), b) transition to turbulence starts 

earlier at station 1 (max subgrid TKE), and c) the location of the peak of u' is away from the 

wall indicating the presence of a bubble. 

5.3. Re - 50,000 steady VGJs cases 

In the case of Re=50,000, and B =0.5, the flow was still separated and not attached. A 

value of B=2 was needed to get the flow attached with very small separation near the trailing 

edge. Below are more detailed results for the Re=50,000 case. 

Pressure coefficient plotted versus dimensionless distance along the suction side of an 

airfoil is presented in Fig. 8. Computational LES results for two blowing ratios (B = 0.5 and B 

= 2) are compared to the experiment with the same blowing conditions. For B = 0.5 large 

separation is shown by both LES and experiment, which is indicated by the large plateau in 

Cp downstream of the suction peak (starting at s/Ls = 0.5 - upstream of station 1). For B = 2 

separation is reduced in size and delayed to a further downstream location, compared to the B 

= 0.5 case. The location of the start of the plateau in Cp has moved downstream to s/Ls = 0.8 

(between stations 4 and 5). The LES is well below the experiment and the experiment does 

not show any plateau. The Transition-sst model over predicts Cp downstream of the s/Ls = 

0.6 (starting at station 2), compared to LES and experiment for the B = 2 case. This indicates 

a larger separation bubble predicted by the Transition-sst model with the location of the 

separation moved upstream. 

Fig. 9 shows velocity profiles normalized by the nominal exit velocity, plotted versus 

dimensionless distance from the wall in the direction normal to the wall. Computational 

velocity profiles from LES for B = 0.5 and B = 2 are presented. Experimental and Transitionsst model results are only available for B = 2. The results for 6 measurement stations from 

LES for B = 0.5, show separation present already at station 1 and continuous through station 

6, which is consistent with the location of the plateau in Cp, observed from Fig. 8. For B = 2 

LES shows separation started between stations 4 and 5 with no reattachment. Separation 

bubble is smaller than for B = 0.5. Results are in a reasonably good agreement with 

experiment with some underprediction of the velocities at stations 3-6. The experiment shows 

low velocities near the wall, indicating that it may be approaching separation near the trailing 

edge, but the boundary layer does not appear to separate. The Transition-sst model 

underpredicts velocities near the wall and shows larger size of the bubble, compared to LES 

and experiment. This model predicts starting separation earlier i.e. at station 2. 

NASA/CR—2012-217415 83
8 

 RMS u' profiles normalized by the nominal exit velocity are plotted versus 

dimensionless distance from the wall in Fig. 10. Computational (LES) and experimental 

profiles are for B = 2. CFD overpredicts u' at stations 3-6. The location and the magnitude of 

the maximum u' are in a reasonable agreement with experiment.

In summary, for B = 0.5: a) separation starts already at station 1, and b) transition to 

turbulence starts between stations 3 and 4 (max subgird TKE, u' and X-vorticity locations). 

For B = 2: a) separation starts near station 3 (Figs. 8, and 9), and b) transition to turbulence 

starts before or at station 1 (max subgrid TKE), and c) The location of the peak of u' is away 

from the wall indicating the presence of a small bubble. 

5.4. Re - 100,000 steady VGJs cases 

In the case of Re=100,000 and B=0.25 the flow separates with reattachment 

downstream the suction peak. At higher blowing ratios (B=1) the flow becomes attached 

along the whole airfoil length. Below are more detailed results for the Re=100,000 case. 

On the pressure coefficient plot (Fig. 11) computational (LES) results for B = 0.25 and 

1 are compared to experimental data for the same blowing ratios. Blowing with B = 0.25 

causes the flow to reattach after a small separation region starting near s/Ls = 0.6 in both CFD 

and experiment. In the higher blowing ratio case (B=1) experimental and computational Cp 

profiles show attached flow at all locations on the suction side of the airfoil. The 

computational Cp profile from the Transition-sst model is shown only for B=1. The 

Transition-sst model predicts separated flow on the suction side of the airfoil, starting at s/Ls 

= 0.5 with no reattachment downstream. Overall LES is in a good agreement with 

experiment, while RANS is not able to predict the flow correctly.

Fig.12 shows velocity profiles from LES for B = 0.25 and B = 1 and experimental 

velocity profiles for B = 0.75. The Transition-sst model results are shown for B = 1. The 

results for 6 measurement stations from LES for B = 0.25 show separation starting at station 2 

and reattached flow at station 4, which is consistent with the location of the plateau in Cp, 

observed from Fig. 11. LES with B = 1 shows attached flow at all six measurement stations. 

Experimental results for B = 0.75 indicate attached flow as well. For stations 1-4 LES 

velocity profiles for B=1 match the experimental profiles for B=0.75 very well. At stations 5 

and 6 the CFD results are in reasonable agreement with the experimental data for B=0.75.

The Transition-sst model predicts separation starting at station 2 and continuing through 

station 6. Velocity profiles from this model are significantly underpredicted compared to LES 

and experiment. 

RMS of the fluctuating component of streamwise velocity normalized by the nominal 

exit velocity is plotted versus dimensionless distance from the wall in the direction normal to 

the wall in Fig.13. The results for 6 measurement stations located downstream of the suction 

peak of the airfoil are shown for B = 1 (CFD) and B = 0.75 (experiment). The agreement 

between CFD and the experiment is good, including the location of the maximum u'. 

 In summary for B = 0.25, separation starts at station 2, flow reattachment is observed 

at station 4. For B = 1: a) attached flow is observed on the whole length of the airfoil's suction 

side (see Figs. 11 and 12), b) transition to turbulence starts at station 3 and c) the location of 

the peak of u' is close to the wall indicating attached flow.

6. Summary and Conclusions

This paper describes active flow control utilizing steady VGJs both experimentally 

and computationally for the L1A airfoil. This study is based on previous work by the authors, 

both experimentally (on a linear cascade in a low speed wind tunnel) and computationally 

(URANS) for the baseline cases (without VGJs). This study is enhanced by additional grid 

NASA/CR—2012-217415 84
9 

independence study for the necessary grid resolution around the jets and in the spanwise 

direction. URANS approach is compared with LES to test the ability to accurately predict 

effect of VGJs on the boundary layer separation. 

Cases were considered at Reynolds numbers (based on the suction surface length and 

the nominal exit velocity from the cascade) of 25,000, 50,000 and 100,000. In all cases 

without flow control, the boundary layer separated and did not reattach. 

6.1. Re=25,000 

Experimental data for Cp show that for B=1, the flow separates at s/Ls between 0.53 

and 0.59, with no reattachment; while at B=3, it is not clear if the flow is attached along the 

airfoil. The velocity profiles data, however, show that the flow separates between s/Ls =0.59 

and 0.69. Furthermore the location of the peak of u' is away from the wall indicating the 

presence of a bubble. 

The LES results are in good agreement with the data for B=1; for B=3 both LES and 

Transition-sst model (for both Cp and U/Ue) are in good agreement with data up to s/Ls = 

0.8. The LES data for u' compare well with the experiment. Overall the LES predictions are 

in much better agreement with the data.

6.2. Re=50,000 

Experimental data for both Cp and velocity profiles show that for B=0.5, the flow 

separates at s/Ls = 0.53 with no reattachment; while at B=2, flow is attached along the airfoil. 

The LES results are in excellent agreement with the data for B=0.5 and showing the 

same trend for B=2. Also, for B=2, LES compares reasonably well with data including 

magnitude and location of u'. As for the Transition-sst model for B=2 it overpredicts Cp

downstream of the s/Ls = 0.6, and underpredicts velocities near the wall. This indicates a 

larger separation bubble predicted by the Transition-sst model with the location of the 

separation moved upstream. Overall the LES predictions are far superior to the Transition-sst 

model and in much better agreement with the data. 

6.3. Re=100,000 

 Experimental data for Cp show that for B=0.25, the flow separates at s/Ls = 0.53 with 

reattachment at s/Ls = 0.7; while at B=1, the flow is fully attached along the airfoil. Also, the 

velocity profiles data for B=0.75 show that the flow does not separate. Furthermore the 

location of the peak of u' is close to the wall indicating attached flow. 

The LES results are in excellent agreement with the data. On the other hand the 

Transition-sst model, with B=1, shows separation at s/Ls = 0.59 and no reattachment; it 

significantly underpredicts velocity profiles compared to LES or experiment. Again, overall 

the LES predictions are far superior to the Transition-sst model and in much better agreement 

with the data.

7. Acknowledgements 

 This work was sponsored by the National Aeronautics and Space Administration. The grant 

monitor is Dr. Anthony Strazisar of the NASA Glenn Research Center. The support of the 

United States Naval Academy Technical Support Department Shop and Fluids Laboratory is 

greatly appreciated. We greatly appreciate the computer time provided for us by the Ohio 

Super Computer (OSC). The OSC Computer Cluster has been made available as part of the 

Center’s mission to support Ohio Universities. 

NASA/CR—2012-217415 85
10

8. References

1. J.P. Johnston and M. Nishi. Vortex Generator Jets. Means for Flow Separation Control. 

AIAA Journal, 28, pp. 989-994, 1990. 

2. R.J. Volino, O. Kartuzova, and M.B. Ibrahim. Experimental and Computational 

Investigations of Low-Pressure Turbine Separation Control using Vortex Generator Jets. 

GT2009-59983, Proceedings of ASME Turbo Expo 2009: Power for Land, Sea and Air 

GT2009, June 8-12, 2009, Orlando, Florida, USA. 

3. D.P. Rizzetta and M.R. Visbal. Numerical Simulation of Separation Control for 

Transitional Highly Loaded Low-Pressure Turbines. AIAA Journal, 43, pp.1958-1967, 

2005. 

4. D. Postl, A. Gross, H.F. Fasel. Numerical Investigation of Active Control for LowPressure Turbine Blade Separation. AIAA2004-750, Proceedings of 42nd AIAA 

Aerospace Science Meeting and Exhibit, January 5-8, 2004, Reno, Nevada, USA, 2004. 

5. V.K. Garg. Low-Pressure Turbine Separation Control - Comparison with Experimental 

Data. ASME Paper GT-2002-30229, 2002. 

6. F.R. Menter. Two-Equation Eddy-Viscosity Turbulence Models for Engineering 

Applications. AIAA Journal, 32, pp. 1598-1605, 1994. 

7. J.P. Clark. Private Communication, Air Force Research Laboratory, 2007. 

8. X.F. Zhang and H. Hodson. Combined Effects of Surface Trips and Unsteady Wakes on 

the Boundary Layer Development of an Ultra-High-Lift LP Turbine Blade. ASME 

Journal of Turbomachinery, 127, pp. 479-488, 2005. 

9. J.P. Bons, J. Pluim, K. Gompertz and M. Bloxham. The Application of Flow Control to 

an Aft-Loaded Low Pressure Turbine Cascade with Unsteady Wakes. ASME Paper 

GT2008-50864, 2008. 

10. M.B. Ibrahim, O. Kartuzova, and R.J. Volino. Experimental and Computational 

Investigations of Separation and Transition on a Highly Loaded Low Pressure Turbine 

Airfoil: Part 1 – Low Freestream Turbulence Intensity. ASME Paper IMECE2008-68879, 

2008. 

11. R.J. Volino, O. Kartuzova, and M.B. Ibrahim. Experimental and Computational 

Investigations of Separation and Transition on a Highly Loaded Low Pressure Turbine 

Airfoil: Part 2 – High Freestream Turbulence Intensity. ASME Paper IMECE2008-

68776, 2008. 

12. Fluent, Inc., 2005. Fluent 6.3 - User Guide.

13. F.R. Menter, R.B. Langtry, S.R. Likki, Y.B. Suzen, P.G. Huang and S. Völker. A 

Correlation Based Transition Model Using Local Variables – Part I: Model Formulation. 

ASME Journal of Turbomachinery, 128, pp. 413-422, 2006.

14. W.W. Kim and S. Menon. Application of the localized dynamic subgrid-scale model to 

turbulent wall-bounded flows. Technical Report AIAA-97-0210, American Institute of 

Aeronautics and Astronautics, 35th Aerospace Sciences Meeting, Reno, NV, 1997. 

15. S.E. Kim. Large eddy simulation using unstructured meshes and dynamic subgridscale 

turbulence models. Technical Report AIAA-2004-2548, American Institute of Aeronautics 

and Astronautics, 34th Fluid Dynamics Conference and Exhibit, 2004. 

16. U. Piomelli and J.R. Chasnov. Large-Eddy Simulations: Theory and Applications. 

Turbulence and Transition Modeling, lecture notes from the ERCOFTAC/IUTAM

summer school held in Stockholm, 12-20 June, 1995. 

NASA/CR—2012-217415 86
11

Figure 1: Computational grid in the jet vicinity

Figure 2: Pressure coefficient, Re = 25,000 

Figure 3: Velocity profiles, Re = 25,000 

 

Figure 4: RMS of the fluctuating 

component of the streamwise velocity, Re 

= 25,000 

Figure 5: Instantaneous isosurfaces of 

Vx=0.01 m/s, Re = 25,000

Figure 6: Instantaneous X-vorticity, 

Re = 25,000

Figure 7: Subgrid turbulence kintic energy 

(TKE) at the spanwise oriented planes of 6 

measurement stations , Re = 25,000 

7a) B =1 

7b) B =3

NASA/CR—2012-217415 87
12

Figure 8: Pressure coefficient, Re = 50,000 

Figure 9: Velocity profiles, Re = 50,000 

Figure 10: RMS of the fluctuating 

component of the streamwise velocity, 

Re = 50,000 

Figure 11: Pressure coefficient, Re = 100,000 

Figure 12: Velocity profiles, Re = 100,000 

Figure 13: RMS of the fluctuating 

component of the streamwise velocity, 

Re = 100,000 

NASA/CR—2012-217415 88
ASME-ATI-UIT 2010 Conference on Thermal and Environmental Issues in Energy Systems 

16 – 19 May, 2010, Sorrento, Italy 

INTRODUCTION 

Gas turbine engines power nearly all commercial and military 

aircraft and are used extensively for ship propulsion and land based 

power generation. Given the hundreds of billions of liters of fuel 

used in gas turbines each year, any improvement in their efficiency, 

even an improvement of only 1%, would result in a huge savings in 

fuel and reduction in emissions. The low-pressure turbine (LPT) is 

the heaviest component in a gas turbine engine and produces the net 

power from the engine. Increases in LPT component efficiency 

result in almost equal increase in overall engine efficiency [1]. One 

way to potentially improve the LPT is to reduce part count, weight 

and cost through the use of very highly loaded blading. The 

limitation on loading is boundary layer separation, which leads to 

partial loss of lift and higher aerodynamic losses (e.g. Mayle [2]). 

Separation and separated flow transition, which can lead to 

boundary layer reattachment, have received considerable attention, 

as noted by Volino [3,4]. In general, previous work shows that the 

strong acceleration on the leading section of the airfoil keeps the 

boundary layer thin and laminar. When separation does occur, it is 

usually just downstream of the suction peak. If transition then 

occurs in the shear layer over the separation bubble, it is typically 

rapid and often causes the boundary layer to reattach [4,5]. 

Airfoils can be designed with high resistance to separation, as 

described by Praisner and Clark [6], but a loading limit will always 

exist, above which separation will still occur. If flow control were 

incorporated in the design of an advanced airfoil, it might be 

possible to increase the loading limit. Passive devices such as 

boundary layer trips (eg. Zhang et al. [7] and Bohl and Volino [8]) 

have been shown effective for separation control and have the 

distinct advantage of simplicity, but they also introduce parasitic 

losses and cannot be adjusted to account for changes in flow 

conditions. Active devices would be more costly and potentially 

risky in terms of reliability, but could potentially provide better 

control over the entire operating range of interest. 

The literature contains many examples of active separation 

control. In turbomachinery, plasma devices, as used by Huang et al. 

[9], could be viable. Vortex generator jets (VGJs), as introduced by 

Johnston and Nishi [10], are another alternative and the subject of 

the present study. Blowing from small, compound angled holes is 

used to create streamwise vortices. The vortices bring high 

momentum fluid into the near wall region, which can help to control 

separation. The most effective VGJs enter the boundary layer at a 

relatively shallow pitch angle (typically 30 to 45 degrees) relative to 

the wall and a high skew angle (45 to 90 degrees) relative to the 

main flow. The jets also promote transition, and turbulent mixing 

helps to mitigate separation. Several studies (e.g. [11,12]) used 

VGJs on the highly loaded Pack B LPT airfoil. Separation was 

essentially eliminated, even at the lowest Reynolds number 

considered, (Re=25000 based on suction surface length and nominal 

exit velocity). Pulsed jets were more effective than steady jets. The 

initial disturbance created by each pulse caused the boundary layer 

to attach. The turbulence was followed by a calmed period 

(Gostelow et al. [13] and Schulte and Hodson [14]) during which the 

boundary layer was very resistant to separation but very laminar-like 

in terms of its fluctuation levels and low losses. When the time 

between pulses was long enough, the boundary layer did eventually 

relax to a separated state, but due to the control which persisted 

during the calmed period, the VGJs were effective even with low jet 

pulsing frequencies, duty cycles and mass flow rates. Since the 

boundary layer was attached and undisturbed for much of the jet 

pulsing cycle, profile losses were low. 

In the present study, the very highly loaded L1A airfoil is used. 

The L1A is an aft loaded blade designed at the Air Force Research 

Laboratory (AFRL) and available on a limited basis from Clark [15]. 

Dimensions as used in the present study are given in Table 1. The 

L1A has a Zweifel coefficient of 1.35, which corresponds to 10% 

higher loading than the “ultra-high lift” airfoils described by Zhang 

and Hodson [7], and 17% higher loading than the Pack B. Because 

the L1A is highly loaded and aft loaded, it is prone to separation, as 

documented in Bons et al. [16] and Volino [3]. In cases without 

flow control and low Reynolds numbers, the boundary layer 

separates and does not reattach, in spite of transition to turbulence in 

the shear layer over the separation bubble. This result contrasts with 

the results of studies on less aggressive airfoils, which all showed 

reattachment after transition. The failure to reattach results in a 20% 

loss in lift and increases profile losses by up to a factor of 7 

compared to high Reynolds number cases. 

SEPARATION CONTROL ON HIGH LIFT LOW PRESSURE TURBINE AIRFOILS 

USING PULSED JET VORTEX GENERATOR JETS 

Ralph J. Volino*, Mounir B. Ibrahim° 

*United States Naval Academy, Annapolis, MD 21402 USA 

°Cleveland State University, Cleveland, OH 44115 USA 

ABSTRACT 

Boundary layer separation control has been studied using vortex generator jets (VGJs) on a very high lift, low-pressure turbine

airfoil. Experiments were done under low (0.6%) freestream turbulence conditions on a linear cascade in a low speed wind 

tunnel. Pressure surveys on the airfoil surface and downstream total pressure loss surveys were documented. Instantaneous 

velocity profile measurements were acquired in the suction surface boundary layer. Cases were considered at Reynolds 

numbers (based on the suction surface length and the nominal exit velocity from the cascade) of 25000 and 50000. Jet pulsing 

frequency and duty cycle were varied. In cases without flow control, the boundary layer separated and did not reattach. With 

the VGJs, separation control was achieved. At sufficiently high pulsing frequencies, separation control was possible with low 

jet velocities and 10% duty cycle. At lower frequencies, a 50% duty cycle helped by separating the disturbances associated 

with the jets turning on and turning off, thereby doubling the frequency of separation control events above the pulsing 

frequency. Phase averaged velocity profiles and wavelet spectra of the velocity show the VGJ disturbance causes the 

boundary layer to reattach, but that it can re-separate between disturbances. When the disturbances occur at high enough 

frequency, the time available for separation is reduced, and the separation bubble remains closed at all times. 

NASA/CR—2012-217415 89
Separation control with VGJs has been demonstrated on the L1A 

airfoil by Bons et al. [16], who considered a case with Re=50000, 

background freestream turbulence TI=3%, and periodic wakes, and 

by Volino et al. [17], who considered cases with TI=0.6% and 

documented pressure distributions on the airfoils and total pressure 

losses. Cases were considered at Reynolds numbers from 25000 to 

100000 (10000 to 40000 based on inlet velocity and axial chord). 

Jet pulsing frequency was varied from F=0.14 to 1.12 with duty 

cycles of 10% and 50% and blowing ratios ranging from 0.25 to 3.0. 

In agreement with previous studies, pulsed jets were more effective 

than steady jets. Separation control was achieved at Re=25000 with 

B=0.75, F=0.56 and D=10%, and at Re=50000 with F=0.56, B=0.25 

and D=10%. Partial control was possible at lower frequencies, 

particularly with higher blowing ratios or duty cycles. Lower 

frequencies were more effective at Re=50000 than at Re=25000. 

Effective separation control resulted in a 20% increase in lift, and up 

to a 60% reduction in total pressure loss, dropping from about 5 

times the high Re value to about twice the high Re value. 

In the present study, the cases of Volino et al. [17] are examined 

in more detail. In addition to the measurements presented 

previously, velocity in the suction surface boundary layer is used to 

document and explain the flow separation and reattachment, and the 

effect of the VGJs. 

EXPERIMENTAL FACILITY AND MEASUREMENTS

Experiments were conducted in a closed loop wind tunnel with a 

seven blade linear cascade located in the wind tunnel’s third turn, as 

shown in Fig. 1a. A fine screen located upstream of the cascade is 

used to provide uniform inlet conditions and 0.6% freestream 

turbulence intensity as noted in Volino [3]. The freestream 

turbulence intensity in an engine is expected to be of the order 4%, 

which would result in more rapid transition of the boundary layer 

and resistance to separation, as shown in Volino et al. [18]. The 

effect of wakes from upstream airfoils would be similar. The present 

experiments provide a more challenging case for flow control and a 

baseline for cases with higher freestream turbulence and wakes, 

some of which are documented in Volino et al. [19]. 

A tailboard and flaps, shown in Fig. 1a, are used to produce 

periodicity at high Reynolds numbers. At low Reynolds numbers, 

when significant separation bubbles are present, the periodicity is not 

as good due to suppression of the separation bubble thickness on the 

blades closest to the tailboard as discussed in Volino [3]. In cases 

with effective flow control, periodicity is reestablished. 

Each blade in the cascade has a central cavity which extends 

along the entire span. As explained in Volino et al. [17], compressed 

air is supplied to the cavities through fast response solenoid valves. 

A single spanwise row of holes was drilled into the suction surface 

of each blade at the inviscid pressure minimum location, s/Ls=0.5 

(x/Cx=0.62), which is about the optimal location for flow control 

devices. The holes are 0.8 mm in diameter and drilled at 30° to the 

surface and 90° to the main flow direction, as shown in Fig. 1b. The 

hole spacing is 10.6 diameters, and the length to diameter ratio is 12. 

With steady blowing and B=1, the mass flow rate of the jets is 0.04% 

of the main flow. With pulsed jets the mass flow is proportionally 

lower. The solenoid valves pulse the VGJs, and the pulsing 

frequency in dimensionless form is F=fLj-te/Uave. Blowing ratio for 

pulsed jets is the ratio of maximum velocity in the pulse to local 

freestream velocity, as explained in Volino et al. [17]. 

The center blade, designated B4 in Fig. 1, contains pressure taps 

near the spanwise centerline. Stagnation pressure is measured with a 

pitot tube upstream of the cascade. The uncertainty in the suction 

side pressure coefficients, Cp, is 0.07. Total pressure losses are 

documented using a Kiel probe traversed across three blade 

spacings, 0.63 Cx downstream of the cascade. 

Velocity profiles on the suction surface were measured near the 

midspan at the six streamwise stations listed in Table 2 with a hotwire probe. At each measurement location, data were acquired for 

26 seconds at a 20 kHz sampling rate (219 samples). Data were 

acquired at 40 wall normal locations in each profile. The data were 

both time averaged and ensemble averaged based on the phase 

within the jet pulsing cycle at 24 dimensionless times, t/T, within the 

pulsing cycle. 

Table 1: Cascade parameters 

Axial 

Chord, Cx

[mm] 

True 

Chord 

[mm] 

Pitch, 

Lφ

[mm] 

Span 

[mm] 

Suction 

side, Ls

[mm] 

Inlet 

flow 

angle 

Exit 

flow 

angle 

134 146 136 724 203 35° 60° 

tailboard

flap

flap

screen

B1

B2

B3

B4

B5

B6

B7

a) 

(b) 

Fig. 1 Drawings of test section: a) linear cascade, b) airfoil 

with VGJ holes and cross section of hole geometry 

Table 2: Velocity profile measurement stations 

Station 1 2 3 4 5 6 

s/Ls 0.53 0.59 0.69 0.78 0.88 0.97 

x/Cx 0.65 0.72 0.80 0.86 0.92 0.97 

Wavelet spectra of the fluctuating velocity were computed using 

the method described in Volino [20]. In contrast to Fourier spectra, 

in which a signal is transformed from the time domain to the 

frequency domain, wavelet spectra provide the frequency content of 

a signal on a time resolved basis. The Mexican Hat wavelet was 

used for the analysis. The wavelet spectra were ensemble averaged 

to show frequency as a function of phase within the pulsing cycle. 

RESULTS 

Re=25000 

Pressure profiles for cases with Re=25000 are shown in Fig. 2. 

The inviscid profile for the L1A airfoil is shown as a reference. The 

low peak followed by a plateau in the case without jets indicates 

separation without reattachment. As shown in Volino et al. [17], 

steady blowing with B=2.0 or lower has no effect on separation, but 

with B=3.0, there are some signs of reattachment, although the Cp

profile remains significantly different from the inviscid profile. 

With B=1.0, pulsed jets with F=0.28 and D=10% have little effect, 

but with F=0.28 and D=50% the Cp values drop near the trailing 

edge, indicating some reattachment after a large separation bubble. 

The F=0.56 and F=1.12 profiles with D=10% are very similar to the 

F=0.28, D=50% case. Total pressure loss profiles are shown in Fig. 

2b and agree with the Cp profiles. A high Reynolds number 

(300000) case from Volino [3] is shown for comparison. The loss, 

ψ, is shown as a function of distance across the cascade, φ, 

normalized on the blade spacing Lφ. The origin, φ=0, corresponds to 

the location downstream of the trailing edge of the center blade (B4 

in Fig. 1a) in the flow direction. Steady blowing with B=3.0 reduces 

NASA/CR—2012-217415 90
0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

Steady,B=3.0

F=0.28,D=10,B=1.0

F=0.56,D=10,B=1.0

F=1.12,D=10,B=1.0

F=0.28,D=50,B=1.0

Inviscid

−1.5 −1 −0.5 0 0.5 1 1.5 0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4

B5 B6

No jets

Steady,B=3.0

F=0.28,D=10,B=1.0

F=0.56,D=10,B=1.0

F=1.12,D=10,B=1.0

High Re

Fig. 2 Pressure results for Re=25000 cases: a) Cp, b) total 

pressure loss 

the loss peaks somewhat, reestablishes periodicity, and causes the 

loss peaks to shift to the right. The shift indicates an increase in 

flow turning of about 6°. The integrated total pressure loss for the 

passage, however, actually increases by 28% over the no-jet case due 

to the losses in the jets between the plenum and hole exit, as 

documented in Volino et al. [17]. Pulsed jets at F=0.28, D=10% and 

B=1.0 have no effect in reducing losses. The F=0.56 and 1.12 cases 

with D=10% are very similar to each other and have about 29% 

lower losses than the no-jet case. The flow turning is increased by 

about 9° compared to case without jets. In these higher F cases the 

integrated loss is still almost 4 times higher and the flow turning is 

4° lower than the high Re case. The exit flow angle for the high Re

case is within 1° of the design exit flow angle. 

Figure 3 shows time averaged velocity profiles. The top row 

shows the mean velocity at the six streamwise stations of Table 2, 

and the lower row shows the rms fluctuating streamwise velocity, u′. 

Without flow control, the boundary layer has separated by Station 1 

and the separation bubble grows at the downstream stations. The 

peak in u′ is in the shear layer far from the wall. With steady B=3.0, 

u′ is higher at the upstream stations, and separation appears to be 

delayed until Station 2. The boundary layer still separates, however, 

and does not reattach. The separation bubble is about 3⁄4 the 

thickness of the no-jet case. The reduction in bubble thickness 

agrees with the slightly lower loss peaks and somewhat greater flow 

turning in Fig. 2 compared to the no-jet case. With F=0.28 and 

D=10%, u′ is higher and the shear layer slightly thicker than in the 

no-jet case, but there is no significant change in the separation 

bubble thickness, in agreement with the Fig. 2 results. With F=0.28 

and D=50% and with F=0.56 or 1.12 and D=10%, a separation 

bubble is still visible between Stations 2 and 5, but it is much thinner 

than in the other cases, and the boundary layer is reattached by the 

trailing edge. The peak in u′ is close to the wall. 

Phase averaged mean velocity profiles for the F=0.28 cases are 

shown in Fig. 4. The columns correspond to the six streamwise 

stations, and rows are for different phases in the pulsing cycle. With 

D=10%, the boundary layer separates and does not reattach at most 

phases. The low, nearly constant velocity near the wall indicates the 

bubble. The measured velocity is not zero in the bubble because the 

hot-wire cannot distinguish direction, so reversed flow and 

0 1 0 1 0 1 0 1 0 1 0 1 0

0.1

0.2

U/Ue

y/Ls

St 12 3 4 5 6

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0

0.1

0.2

u′/Ue

y/Ls

No jets Steady B=3.0 F=0.28 B=1.0 F=0.56 B=1.0 F=1.12 B=1.0 F=0.28 B=1.0

Fig. 3 Time averaged velocity profiles for Re=25000 cases with 

steady jets or D=10% (solid lines), and D=50% (dashed lines): 

top – mean velocity, bottom – rms velocity 

0

0.1

0.083

t/T St. 1 2 3 4 5 6

0.083

t/T St. 1 2 3 4 5 6

0

0.1

0.167

0

0.1

0.25

0

0.1

0.333

0

0.1

0.417

0

0.1

0.5

y/Ls

0

0.1

0.583

0

0.1

0.667

0

0.1

0.75

0

0.1

0.833

0

0.1

0.917

0 1 0 1 0 1 0 1 0 1 0 1 0

0.1

1

U/Ue

Fig. 4 Phase averaged mean velocity profiles for Re=25000, 

F=0.28, B=1.0 cases, columns for six streamwise stations, rows 

for phases in pulsing cycle: blue – D=10%, red – D=50% 

turbulence cause a false positive mean velocity in the bubble. At 

t/T=0.333, 0.5, 0.583, and 0.667 to 0.75 at Stations 3-6 respectively, 

the near wall region of low velocity is less apparent and the velocity 

goes more continuously toward zero at the wall. This indicates 

reattachment at these phases. Figure 5a shows wavelet spectra for 

the F=0.28, D=10% case. The six plots in the figure correspond to 

the six streamwise stations. In each plot, the horizontal axis shows 

dimensionless frequency, fLs/Ue on a log scale, and the vertical axis 

shows dimensionless time, t/T, for one pulsing cycle. Power spectral 

density is computed from instantaneous velocity data at all y

locations and is shown for the y location corresponding to maximum 

time averaged u′ at each station. The contours show the power 

spectral density premultiplied by frequency and normalized by Ue

2

. 

The color scale is the same for all plots. The VGJs creates a 

disturbance at the beginning of a pulse, and the leading edge of this 

disturbance is visible as high contours centered at t/T=0.11, 0.24, 

0.26, 0.36, 0.43 and 0.49 at Stations 1-6 respectively. The arrival 

times at each station indicate that the leading edge of the disturbance 

convects along the surface at about 0.5 times the local freestream 

velocity (marked by solid white line). A second peak appears 

centered at t/T=0.19, 0.39, 0.46, 0.60, 0.75 and 0.91 at Stations 1-6 

NASA/CR—2012-217415 91
t/T

0.1 1 10 0

0.2

0.4

0.6

0.8

1

0.1 1 10 f⋅Ls

/Ue

0.1 1 10 0.1 1 10 0.1 1 10 0.1 1 10

(a)

t/T

0.1 1 10 0

0.2

0.4

0.6

0.8

1

0.1 1 10 f⋅Ls

/Ue

0.1 1 10 0.1 1 10 0.1 1 10 0.1 1 10

(b)

t/T

0.1 1 10 0

0.2

0.4

0.6

0.8

1

0.1 1 10 f⋅Ls

/Ue

0.1 1 10 0.1 1 10 0.1 1 10 0.1 1 10

(c) 

Fig. 5 Wavelet spectra for Re=25000, B=1.0 computed at y 

locations of maximum u′ in time averaged profiles and shown as 

function of time and frequency at six streamwise stations: a) 

F=0.28, D=10%, b) F=0.28, D=50%, c) F=0.56, D=10%

respectively. These peaks are believed to result when the trailing 

edge of the VGJ disturbance passes, and the times indicate a 

convection speed of about 0.3 times the freestream velocity (marked 

by dashed white line). The arrival times of the disturbance in Fig. 5 

are just prior to the reattachment times in Fig. 4. With the strong 

adverse pressure gradient, the boundary layer separates again even 

before the trailing edge of the disturbance passes. Peaks appear in 

Fig. 5 at the dimensionless frequency, fLs/Ue=1.4, and tails of these 

peaks extend to the pulsing frequency, which is 0.7 for this case. A 

higher frequency peak at fLs/Ue≈7 is also visible, particularly at 

Stations 1 and 2. The higher frequency is likely related to shear 

layer transition, and matches the frequency peak observed in 

transition without flow control in Volino et al. [3]. At the 

downstream stations there is a broader range of frequencies for the 

full cycle, since the shear layer has transitioned to turbulent. 

When the duty cycle is increased to 50% with F=0.28, the 

separation bubble is thinner. Figure 4 shows boundary layer 

reattachment at about the same phases as in the D=10% case. It then 

starts to re-separates, but at Stations 5 and 6 it reattaches again about 

half a cycle later. This second reattachment means less time for the 

separation bubble to grow, resulting in a thinner separation at all 

times. The wavelet spectra of Fig. 5b show the disturbance 

responsible for the second reattachment. When the VGJ turns on, it 

creates a disturbance much like that in the D=10% case. When the 

jet turns off half a cycle later, it creates a second disturbance which 

moves down the surface at about the same velocity as the first 

disturbance (marked by magenta lines). This second disturbance 

also causes reattachment. In the D=10% case, the jets turn off only 

0.1 cycle after they turn on, so the on and off disturbances act 

essentially as a single event. With D=50% there is more time 

between the on and off disturbances, so they act as two separate 

events. The result is two effective disturbances per cycle, thereby 

doubling the disturbance frequency above the pulsing frequency and 

providing better separation control. 

Figures 5c shows wavelet spectra for the F=0.56, D=10% case. 

The convection velocity of the disturbances is about the same as 

with F=0.28, but the period, T, is shorter, so the disturbances occupy 

a larger fraction of the cycle, and there is less time between 

disturbances for the separation bubble to grow. The time averaged 

profiles for the F=0.28, D=50% and F=0.56, D=10% cases in Fig. 3 

are very similar, since these two cases have the same effective 

disturbance frequency (due to the doubling of the pulsing frequency 

with D=50% discussed above). The frequencies of the largest peaks 

in Fig. 5c are about fLs/Ue=2.6, which is which is roughly double 

that in the F=0.28 cases, probably due to the higher pulsing 

frequency. As in the F=0.28 case, peaks associated with shear layer 

transition at fLs/Ue≈7 are visible at Stations 1 and 2. The wavelet 

spectra for the F=1.12, D=10% case (not shown) are very similar to 

those with F=0.56, although the peak frequency is increased to about 

fLs/Ue=3.7 since the pulsing frequency is higher. Upstream peaks at 

fLs/Ue≈7 are still visible. Since the disturbance occupy nearly the 

entire cycle by the downstream stations with F=0.56, increasing to 

F=1.12 has little additional effect, as shown by the mean profiles in 

Fig. 3. 

Re=50000 

Pressure profiles for the cases with Re=50000 are shown in Fig. 

6. Without flow control the boundary layer does not reattach. As 

shown in Volino et al. [17], steady jets become effective when 

B≥1.5, and by B=2 the separation is nearly eliminated and the loss 

peak magnitude and width are greatly reduced. The integrated loss 

is about the same as in the no-jet case due to the loss associated with 

the pressure drop from the plenum to the jet exit. Flow turning is 

increased 12° compared to the no-jet case. With pulsed jets and 

B=1.0, pulsing with F=0.14 has only a limited affect with D=10%, 

but with F=0.14 and D=50% the boundary layer reattaches after a 

large separation bubble. Losses drop to about half the value in the 

no-jet case and flow turning increases 12° from the no-jet case. 

With F=0.28 and D=10%, the Cp profile is about the same as in the 

F=0.14, D=50% case, but the loss peak is somewhat wider and the 

shift in the peak indicates about 3° less flow turning. With F=0.56, 

Cp shows a slightly smaller separation bubble than at the lower 

frequencies, the losses are reduced to 40% of the no-jet case, and 

flow turning is 13° higher than the no-jet case. Increasing the duty 

cycle from 10% to 50% has little effect when F=0.56. The losses 

with F=0.56 are about 2.4 times those in the high Re reference case 

and there is about 3° less flow turning due to the presence of the 

separation bubble and thicker boundary layer at Re=50000. 

Figure 7 shows time averaged velocity profiles for the steady and 

D=10%, Re=50000 cases. A large separation bubble without 

reattachment is present without flow control. Steady jets with B=2.0 

eliminate the separation, in agreement with the Cp profiles of Fig. 6. 

With F=0.14 the separation is nearly as large as in the no-jet case, 

but the shear layer is somewhat thicker and u′ levels are higher. 

With F=0.28 there is small separation bubble at Stations 3 and 4, but 

the boundary layer reattaches downstream. With F=0.56 the 

boundary layer appears attached at all stations. Figure 8 shows the 

effect of duty cycle. With F=0.14, increasing the duty cycle from 

10% to 50% greatly reduces the separation bubble thickness, 

although a small bubble is still present when D=50%. When 

F=0.56, duty cycle appears to have no effect, in agreement with the 

pressure results of Fig. 6. The boundary layer is attached with both 

D=10% and 50%. 

Phase averaged mean velocity is shown for the F=0.14 cases in 

Fig. 9. With D=10%, the boundary layer appears to be close to 

reattaching at t/T=0.25, 0.333, 0.417 and 0.5 at Stations 3 through 6 

respectively. With D=50%, reattachment occurs at about the same 

phases as in the D=10% case. The boundary layer then starts to reseparate, but at Stations 5 and 6 it reattaches again about half a cycle 

NASA/CR—2012-217415 92
0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

Steady,B=2.0

F=0.14,D=10,B=1.0

F=0.28,D=10,B=1.0

F=0.56,D=10,B=1.0

F=0.14,D=50,B=1.0

F=0.56,D=50,B=1.0

Inviscid

−1.5 −1 −0.5 0 0.5 1 1.5 0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4

B5 B6

No jets

Steady,B=2.0

F=0.14,D=10,B=1.0

F=0.28,D=10,B=1.0

F=0.56,D=10,B=1.0

F=0.14,D=50,B=1.0

F=0.56,D=50,B=1.0

High Re

Fig. 6 Pressure results for Re=50000 cases: a) Cp, b) total 

pressure loss

later. This second reattachment means less time for the separation 

bubble to grow, resulting in a thinner separation at all times. The 

first reattachment corresponds to the arrival of the disturbance 

created when the VGJs are turned on. The second reattachment 

corresponds to the disturbance created when the VGJs turn off. As 

in the Re=25000 case, increasing the duty cycle to 50% separates in 

time the disturbances created when the VGJs turn on and off, thereby 

doubling the frequency of separation control events above the 

pulsing frequency. The result, as shown in Fig. 9, is a reattached 

boundary layer for most of the cycle by the downstream stations. 

The same effect can be achieved by doubling the pulsing frequency 

to F=0.28 with D=10%. 

CONCLUSIONS 

The effect of vortex generator jets on the flow over the very high 

lift L1A airfoil was studied under low freestream turbulence 

conditions. Reynolds numbers based on suction surface length and 

nominal exit velocity of 25000 and 50000 were considered. Without 

flow control, the boundary layer separated and did not reattach. 

Flow control with VGJs was possible even at Re=25000. In 

agreement with previous studies, pulsed jets were more effective 

than steady jets. Effective separation control resulted in a 20% 

increase in lift and up to a 60% reduction in total pressure loss 

compared to baseline cases at the same Reynolds number. Loss 

values still remain higher than in high Reynolds number cases. 

Phase averaged velocity profiles and wavelet spectra show the 

boundary layer intermittently reattaching as disturbances pass and 

then separating between disturbances. Increasing the pulsing 

frequency reduces the time available for separation. When the time 

available is sufficiently small, the boundary layer remains attached 

at all times. At Re=25000, separation was nearly fully controlled for 

the full pulsing cycle when F=0.5 and D=10%. Higher frequency 

pulsing provided little additional benefit. At Re=50000 the 

separation bubble grows more slowly, so F=0.3 is sufficient. At 

lower pulsing frequencies, increasing the duty cycle to 50% is 

helpful. The most effective disturbances for controlling separation 

are created when the VGJs turn on and off, and increasing the duty 

0 1 0 1 0 1 0 1 0 1 0 1 0

0.1

0.2

U/Ue

y/Ls

St 12 3 4 5 6

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0

0.1

0.2

u′/Ue

y/Ls

No jets Steady B=2.0 F=0.14 B=1.0 F=0.28 B=1.0 F=0.56 B=1.0

Fig. 7 Time averaged velocity profiles at six streamwise 

stations for Re=50000 cases with steady jets or D=10%: top – 

mean velocity, bottom – rms velocity 

0 1 0 1 0 1 0 1 0 1 0 1 0

0.1

0.2

U/Ue

y/Ls

St 12 3 4 5 6

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0

0.1

0.2

u′/Ue

y/Ls

No jets F=0.14 D=10% F=0.56 D=10% F=0.14 D=50% F=0.56 D=50%

Fig. 8 Time averaged velocity profiles at six streamwise stations 

for Re=50000 cases with D=10% (solid lines), and D=50% 

(dashed lines): top – mean velocity, bottom – rms velocity

0

0.1

0.083

t/T St. 1 2 3 4 5 6

0.083

t/T St. 1 2 3 4 5 6

0

0.1

0.167

0

0.1

0.25

0

0.1

0.333

0

0.1

0.417

0

0.1

0.5

y/Ls

0

0.1

0.583

0

0.1

0.667

0

0.1

0.75

0

0.1

0.833

0

0.1

0.917

0 1 0 1 0 1 0 1 0 1 0 1 0

0.1

1

U/Ue

Fig. 9 Phase averaged mean velocity profiles for Re=50000, 

F=0.14, B=1.0 cases, columns for six streamwise stations, rows 

for phases in pulsing cycle: blue – D=10%, red – D=50% 

cycle to 50% separates in time the on and off events, thereby 

doubling the frequency of flow control events above the pulsing 

frequency. 

NASA/CR—2012-217415 93
ACKNOWLEDGMENT 

This work was sponsored by the National Aeronautics and Space 

Administration under grant NNC07IA10I. The grant monitors were 

Drs. Anthony Strazisar and James Heidmann of the NASA Glenn 

Research Center. The support of the United States Naval Academy 

Technical Support Department Shop and Fluids Laboratory is greatly 

appreciated. 

NOMENCLATURE 

Symbol Quantity SI Unit 

B blowing ratio 

Cp ( ) 2

T − ρU/PP2 e , pressure coefficient 

Cx axial chord m 

D duty cycle % 

F fLj-te/Uave

f frequency Hz 

Lj-te distance from VGJs to trailing edge m

Ls suction surface length m 

Lφ blade spacing (pitch) m 

P pressure kPa 

PS upstream static pressure kPa 

PT upstream stagnation pressure kPa 

PTe downstream stagnation pressure kPa 

Re UeLs/ν, exit Reynolds number 

s streamwise coordinate m 

T period of jet pulsing cycle s 

t time s 

Uave average freestream velocity between m/s 

VGJ holes and trailing edge 

Ue nominal exit freestream velocity m/s 

u′ rms fluctuating streamwise velocity m/s 

x axial distance from leading edge m 

φ coordinate along blade spacing m 

ν kinematic viscosity m2/s 

ρ density kg/m3

ψ (PT-PTe)/(PT-PS), total pressure loss 

 

REFERENCES 

[1] Wisler, D.C., 1998, “The Technical and Economic Relevance 

of Understanding Boundary Layer Transition in Gas Turbine 

Engines,” Minnowbrook II, 1997 Workshop on Boundary Layer 

Transition in Turbomachines, LaGraff, J.E., and Ashpis, D.E., 

eds., NASA/CP-1998-206958, pp. 53-64. 

[2] Mayle, R.E., 1991, “The Role of Laminar-Turbulent Transition 

in Gas Turbine Engines,” ASME Journal of Turbomachinery, 

113, pp. 509-537. 

[3] Volino, R.J., 2010, “Separated Flow Measurements on a Highly 

Loaded Low-Pressure Turbine Airfoil,” ASME Journal of 

Turbomachinery, 132, 011007. 

[4] Volino, R.J., 2002, “Separated Flow Transition Under 

Simulated Low-Pressure Turbine Airfoil Conditions: Part 1 – 

Mean Flow and Turbulence Statistics,” ASME Journal of 

Turbomachinery, 124, pp. 645-655. 

[5] Volino, R.J., 2002, “Separated Flow Transition Under 

Simulated Low-Pressure Turbine Airfoil Conditions: Part 2 –

Turbulence Spectra,” ASME Journal of Turbomachinery, 124, 

pp. 656-664. 

[6] Praisner, T.J., and Clark, J.P., 2007, “Predicting Transition in 

Turbomachinery – Part 1: A Review and New Model 

Development,” ASME Journal of Turbomachinery, 129, pp. 1-

13. 

[7] Zhang, X.F., Vera, M., Hodson, H., and Harvey, N., 2007, 

Separation and Transition Control on an Aft-Loaded UltraHigh-Lift LP Turbine Blade at Low Reynolds Numbers: Low-

 

Speed Investigation,” ASME Journal of Turbomachinery, 128, 

pp. 517-527. 

[8] Bohl, D.G., and Volino, R.J., 2006, “Experiments with ThreeDimensional Passive Flow Control Devices on Low-Pressure 

Turbine Airfoils,” ASME Journal of Turbomachinery, 128, pp. 

251-260. 

[9] Huang, J., Corke, T., and Thomas, F., 2003, “Plasma Actuators 

for Separation Control on Low Pressure Turbine Blades,” 

AIAA Paper 2003-1027. 

[10] Johnston, J.P., and Nishi, M., 1990, “Vortex Generator Jets. 

Means for Flow Separation Control,” AIAA Journal, 28, pp. 

989-994. 

[11] Bons, J.P., Sondergaard, R., and Rivir, R.B., 2002, “The Fluid 

Dynamics of LPT Blade Separation Control Using Pulsed Jets,” 

ASME Journal of Turbomachinery, 124, pp. 77-85. 

[12] Volino, R.J., 2003, “Separation Control on Low-Pressure 

Turbine Airfoils Using Synthetic Vortex Generator Jets,” 

ASME Journal of Turbomachinery, 125, pp. 765-777. 

[13] Gostelow, J.P., Walker, G.J., Solomon, W.J., Hong, G., and 

Melwani, N., 1997, “Investigation of the Calmed Region 

Behind a Turbulent Spot,” ASME Journal of Turbomachinery, 

119, pp. 802-809. 

[14] Schulte, V., and Hodson, H.P., 1998, “Prediction of the 

Becalmed Region for LP Turbine Profile Design,” ASME 

Journal of Turbomachinery, 120, pp. 839-846. 

[15] Clark J.P., 2007, Private Communication, Air Force Research 

Laboratory. 

[16] Bons, J.P., Pluim, J., Gompertz, K., Bloxham, M., and Clark, 

J.P., 2008, “The Application of Flow Control to an Aft-Loaded 

Low Pressure Turbine Cascade with Unsteady Wakes,” ASME 

Paper GT2008-50864. 

[17] Volino, R.J., Kartuzova, O., and Ibrahim, M.B., 2009, 

“Experimental and Computational Investigations of LowPressure Turbine Separation Control Using Vortex Generator 

Jets,” ASME Paper GT2009-59983. 

[18] Volino, R.J, Kartuzova, O., and Ibrahim, M., 2008, 

“Experimental and Computational Investigations of Separation 

and Transition on a Highly Loaded Low Pressure Turbine 

Airfoil: Part 2 – High Freestream Turbulence Intensity,” ASME 

Paper IMECE2008-68776. 

[19] Volino, R.J., Kartuzova, O., and Ibrahim, M.B., 2010, 

“Separation Control on a Very High Lift Low Pressure Turbine 

Airfoil Using Pulsed Vortex Generator Jets,” ASME Paper 

GT2010-23567. 

[20] Volino, R.J., 2005, “An Investigation of the Scales in 

Transitional Boundary Layers under High Freestream 

Turbulence Conditions,” Experiments in Fluids, 38, pp. 516-

633. 

NASA/CR—2012-217415 94
1 

Proceedings of ASME Turbo Expo 2010: Power for Land, Sea and Air 

GT2010 

June 14-18, 2010, Glasgow, UK

GT2010-23567 

SEPARATION CONTROL ON A VERY HIGH LIFT LOW PRESSURE TURBINE AIRFOIL 

USING PULSED VORTEX GENERATOR JETS

Ralph J. Volino

Mechanical Engineering Department 

United States Naval Academy 

Annapolis, Maryland 21402-5042 

volino@usna.edu 

Olga Kartuzova and Mounir B. Ibrahim 

Mechanical Engineering Department 

Cleveland State University 

Cleveland, Ohio 44115-2425 

m.ibrahim@csuohio.edu 

ABSTRACT 

 Boundary layer separation control has been studied using vortex 

generator jets (VGJs) on a very high lift, low-pressure turbine airfoil. 

Experiments were done under high (4%) freestream turbulence 

conditions on a linear cascade in a low speed wind tunnel. Pressure 

surveys on the airfoil surface and downstream total pressure loss 

surveys were documented. Instantaneous velocity profile 

measurements were acquired in the suction surface boundary layer. 

Cases were considered at Reynolds numbers (based on the suction 

surface length and the nominal exit velocity from the cascade) of 

25,000 and 50,000. Jet pulsing frequency, duty cycle, and blowing 

ratio were all varied. Computational results from a large eddy 

simulation of one case showed reattachment in agreement with the 

experiment. In cases without flow control, the boundary layer 

separated and did not reattach. With the VGJs, separation control was 

possible even at the lowest Reynolds number. Pulsed VGJs were more 

effective than steady jets. At sufficiently high pulsing frequencies, 

separation control was possible even with low jet velocities and low 

duty cycles. At lower frequencies, higher jet velocity was required, 

particularly at low Reynolds numbers. Effective separation control 

resulted in an increase in lift and a reduction in total pressure losses. 

Phase averaged velocity profiles and wavelet spectra of the velocity 

show the VGJ disturbance causes the boundary layer to reattach, but 

that it can re-separate between disturbances. When the disturbances 

occur at high enough frequency, the time available for separation is 

reduced, and the separation bubble remains closed at all times. 

NOMENCLATURE 

B blowing ratio, maximum jet velocity/local freestream velocity 

Cp ( ) 2

T − ρU/PP2 e , pressure coefficient 

Cpint pressure coefficient integrated in axial direction 

Cpinv Cpint for invisid flow through cascade 

Cx axial chord length 

D duty cycle, fraction of time valve is open 

F fLj-te/Uave, dimensionless frequency 

f frequency 

Lj-te distance from VGJs to trailing edge

Ls suction surface length 

Lφ blade spacing (pitch) 

m j  jet mass flow rate 

m1 main flow mass flow rate 

P pressure 

Pj jet plenum pressure 

PS upstream static pressure 

PT upstream stagnation pressure 

PTe downstream stagnation pressure 

Re UeLs/ν, exit Reynolds number 

s streamwise coordinate, distance from leading edge 

T period of jet pulsing cycle 

t time 

Uave average freestream velocity between VGJs and trailing edge 

Ue nominal exit freestream velocity, based on inviscid solution 

u′ rms fluctuating streamwise velocity 

x axial distance from leading edge 

Δx

+

 grid spacing in axial direction in viscous units 

y+ location of closest grid point to wall in viscous units 

Δz

+ grid spacing in spanwise direction in viscous units 

φ coordinate along blade spacing, normal to axial chord 

ν kinematic viscosity 

ρ density 

ψ (PT-PTe)/(PT-PS), total pressure loss coefficient 

ψint total pressure loss integrated over blade spacing 

INTRODUCTION 

 Boundary layer separation can lead to partial loss of lift and higher 

aerodynamic losses on low-pressure turbine (LPT) airfoils (e.g. 

Hourmouziadis [1], Mayle [2], and Sharma et al. [3]). As designers 

impose higher loading to improve efficiency and lower cost, the 

associated strong adverse pressure gradients on the suction side of the 

airfoil can exacerbate separation problems. In aircraft engines, the 

lower density and therefore lower Reynolds numbers at altitude can 

lead to a component efficiency drop of 2% between takeoff and cruise 

in large commercial transport engines, and possibly as much as 7% in 

smaller engines operating at higher altitudes [4, 5]. Prediction and 

control of suction side separation, without sacrifice of the benefits of 

higher loading is, therefore, crucial for improved engine design. 

NASA/CR—2012-217415 95

This material is declared a work of the U.S. Government and is not subject to copyright protection in the United States. Approved for public release; 

distribution is unlimited. 
2 

 Separation and separated flow transition, which can lead to 

boundary layer reattachment, have received considerable attention. 

Lists of various studies are provided by Volino [6, 7]. In general, 

previous work shows that the strong acceleration on the leading 

section of the airfoil keeps the boundary layer thin and laminar, even 

in the presence of elevated freestream turbulence. When separation 

does occur, it is usually just downstream of the suction peak. If 

transition then occurs in the shear layer over the separation bubble, it 

is typically rapid and often causes the boundary layer to reattach [7, 8]. 

 Airfoils can be designed with high resistance to separation, as 

described by Praisner and Clark [9], but a loading limit will always 

exist, above which separation will still occur. If flow control were 

incorporated in the design of an advanced airfoil, as discussed by Bons 

et al. [10], it might be possible to increase the loading limit. 

Separation control with passive devices such as boundary layer trips 

has been shown effective by Zhang et al. [11], Bohl and Volino [12], 

Volino [13], and others. Passive devices have the distinct advantage 

of simplicity, but they also introduce parasitic losses and cannot be 

adjusted to account for changes in flow conditions. Active devices 

could potentially provide better control over the entire operating range 

of interest and be reduced in strength or turned off to avoid 

unnecessary losses when they are not needed. Potential disadvantages 

of active devices are the cost of implementation and the possibility of 

control failure, particularly in aircraft engines. 

 The literature contains many examples of active separation control. 

A few which could be applied in turbomachinery are discussed in 

Volino [14]. Plasma devices, as used by Huang et al. [15], could be 

viable, and are under active study. Vortex generator jets (VGJs), as 

introduced by Johnston and Nishi [16], are another alternative and the 

subject of the present study. Blowing from small, compound angled 

holes is used to create streamwise vortices. The vortices bring high 

momentum fluid into the near wall region, which can help to control 

separation. The most effective VGJs enter the boundary layer at a 

relatively shallow pitch angle (typically 30 to 45 degrees) relative to 

the wall and a high skew angle (45 to 90 degrees) relative to the main 

flow. Additionally, the jets promote transition, and turbulent mixing 

also helps to mitigate separation. Bons et al. [10] noted that in the 

case of pulsed VGJs, the turbulence effect is more significant than the 

action of the vortices. Bons et al. [4, 17], Volino [14], Volino and 

Bohl [18], McQuilling and Jacob [19], and Eldredge and Bons [20] all 

used VGJs on the highly loaded Pack B LPT airfoil. Separation was 

essentially eliminated, even at the lowest Reynolds number 

considered, (Re=25,000 based on suction surface length and nominal 

exit velocity). Pulsed jets were more effective than steady jets. The 

initial disturbance created by each pulse caused the boundary layer to 

attach. The turbulence was followed by a calmed period (Gostelow et 

al. [21] and Schulte and Hodson [22]) during which the boundary layer 

was very resistant to separation, much like a turbulent boundary layer, 

but very laminar-like in terms of its fluctuation levels and low losses. 

When the time between pulses was long enough, the boundary layer 

did eventually relax to a separated state, but due to the control which 

persisted during the calmed period, the VGJs were effective even with 

low jet pulsing frequencies, duty cycles and mass flow rates. Since the 

boundary layer was attached and undisturbed for much of the jet 

pulsing cycle, profile losses were low. 

 Similar results with pulsed VGJs were found on the L1M airfoil by 

Bons et al. [23]. The L1M is more highly loaded than the Pack B, but 

more resistant to separation because of forward loading. A large 

separation bubble followed by boundary layer reattachment was 

observed at low Reynolds numbers, and pulsed VGJs reduced the size 

of the bubble. 

 In the present study, the very highly loaded L1A airfoil is used. The 

L1A was designed at the Air Force Research Laboratory (AFRL) and 

is available on a limited basis from Clark [24]. It is an aft loaded 

blade with the same flow angles and loading as the L1M. Dimensions 

of the L1A as used in the present study are given in Table 1. Based on 

the design calculations of Clark [24], the L1A has a Zweifel 

coefficient of 1.35, which corresponds to 10% higher loading than the 

“ultra-high lift” airfoils described by Zhang and Hodson [25], and 17% 

higher loading than the Pack B. Because the L1A is aft loaded, it is 

more prone to separation than the L1M, as documented in Bons et al. 

[26], Volino [6], Ibrahim et al. [27], and Volino et al. [28]. In cases 

without flow control and low Reynolds numbers, the boundary layer 

separates and does not reattach, in spite of transition to turbulence in 

the shear layer over the separation bubble. This result contrasts with 

the results of studies on less aggressive airfoils, which all showed 

reattachment after transition. The separation bubble on the L1A is 

about four times thicker than that on the Pack B. The larger distance 

from the shear layer to the wall on the L1A apparently prevents the 

turbulent mixing in the shear layer from reaching the wall and causing 

reattachment. The failure of the boundary layer to reattach results in a 

20% loss in lift and increases profile losses by up to a factor of 7. At 

higher Reynolds numbers the separation bubble is small and the 

boundary layer is attached over most of the airfoil. 

 Separation control with VGJs has been demonstrated on the L1A 

airfoil. Bons et al. [26] considered a case with Re=50,000, background 

freestream turbulence TI=3%, and periodic wakes produced with 

moving rods upstream of the airfoils. The VGJs had a duty cycle 

(fraction of time the jets are on during a pulsing cycle) D=30% and a 

blowing ratio B=2.3 (based on the maximum jet velocity in the pulsing 

cycle and the local freestream velocity). The dimensionless frequency 

of the wake passing and VGJ pulsing were both F=fLj-te/Uave=0.34, 

where Lj-te is the streamwise distance from the VGJ holes to the 

trailing edge, and Uave is the average freestream velocity over this 

distance. Two different streamwise locations for the VGJs were 

considered. The VGJs reduced the separation bubble size and reduced 

total pressure losses. Volino et al. [29] considered cases with 

TI=0.6%. Pressure distributions on the airfoils and total pressure 

losses were documented. Cases were considered at Reynolds numbers 

from 25,000 to 100,000. Jet pulsing frequency was varied from 

F=0.14 to 1.12 with duty cycles of 10% and 50% and blowing ratios 

ranging from 0.25 to 3.0. In agreement with previous studies, pulsed 

jets were more effective than steady jets. At Re=25,000, separation 

control was achieved with B=0.75, F=0.56 and D=10%. Without flow 

control the lift at Re=25,000 was about 80% of the high Re lift, and 

with flow control it was 94% of the high Re lift. Flow control caused 

the total pressure loss to drop from about five times the high Re value 

to about 3 times the high Re value. Once good control was achieved, 

further increasing the frequency, blowing ratio or duty cycle provided 

little additional benefit. Partial control was possible at lower 

frequencies, particularly with higher blowing ratios or duty cycles. At 

Re=50,000 with F=0.56, B=0.25 was sufficient with D=10%. Lower 

frequencies were more effective at Re=50,000 than at Re=25,000. 

Effective separation control resulted in a 20% increase in lift, and up 

to a 60% reduction in total pressure loss, dropping from about 5 times 

the high Re value to about twice the high Re value. 

 In the present study, flow control with vortex generator jets is 

investigated under high (4%) freestream turbulence conditions, which 

is more representative of engine conditions than the low TI considered 

in Volino et al. [29]. Cases with a range of pulsing frequencies and 

blowing ratios are considered. In addition to pressure distributions, 

instantaneous boundary layer velocity measurements are used to 

explain the mechanism by which the VGJs control separation. 

NASA/CR—2012-217415 96
3 

Table 1: Cascade parameters 

Axial 

Chord, Cx

[mm] 

True 

Chord 

[mm] 

Pitch, 

Lφ

[mm] 

Span 

[mm] 

Suction 

side, Ls

[mm] 

Inlet 

flow 

angle 

Exit 

flow 

angle 

134 146 136 724 203 35° 60° 

EXPERIMENTAL FACILITY AND MEASUREMENTS

 Experiments were conducted in a closed loop wind tunnel with a 

seven blade linear cascade located in the wind tunnel’s third turn, as 

shown in Fig. 1a. A coarse grid, consisting of a 1.5 mm thick sheet 

metal plate with 19 mm square holes spaced 25.4 mm apart, center to 

center, in both directions is located upstream of the cascade. In a 

plane perpendicular to the inlet flow and 1.7Cx upstream of the center 

blade, the grid produced uniform flow with TI=6.0% in the streamwise 

component and 4.2% in the cross stream components, for an overall 

intensity of 4.9%. The integral length scale of the freestream 

turbulence is 0.12Cx in the streamwise component and 0.04Cx in the 

other components. The streamwise component was also measured at 

the inlet plane of the cascade in the four center passages, where it had 

decayed to between 4 and 4.2% between blades B2 and B5 and 4.6% 

between blades B5 and B6. Downstream of the cascade, the local TI is 

1.8% across all passages, as documented in Volino et al. [28]. The 

local freestream turbulence intensity in the passage at the beginning of 

the adverse pressure gradient region is 1.4%. The change in TI

through the passage is due mainly to the change in the local freestream 

velocity along with some decay of the turbulence. 

 A tailboard, shown in Fig. 1a, is needed to produce the correct exit 

flow angle from the cascade. Its position was set to produce 

periodicity at high Reynolds numbers. A tailboard on the opposite 

side of the cascade, and inlet guide vanes were found to be 

unnecessary. To produce the correct approach flow to the end blades 

(B1 and B7), the amount of flow escaping around the two ends of the 

cascade is controlled with the flaps shown in Fig. 1a. The inlet flow 

angle was checked with a three-hole pressure probe and found to agree 

with the design angle to within 2° of uncertainty. Good periodicity at 

high Reynolds numbers was obtained in the exit flow. At low 

Reynolds numbers, when significant separation bubbles are present, 

the periodicity is not as good due to suppression of the separation 

bubble thickness on the blades closest to the tailboard. In cases with 

effective flow control, periodicity is reestablished. The lack of 

periodicity in cases with large separation bubbles is considered 

acceptable since the focus of the study is separation control, and not 

documentation of cases with large separation that would be 

unacceptable in practice. This compromise facilitates the study of a 

larger number of cases with flow control by obviating the need to 

adjust the tailboard by trial and error for each case. It also provides for 

better repeatability in the experiments, since the position of the 

tailboard is fixed for all cases. Any improvements made with flow 

control will be larger in practice than documented in the experiment, 

due to the effect of the tailboard in suppressing the bubble size in the 

uncontrolled cases. 

 Each blade in the cascade has a central cavity which extends along 

the entire span. As explained in Volino et al. [29], compressed air is 

supplied to the cavities from a common manifold. The manifold is 

supplied through two fast response solenoid valves (Parker Hannifin 

009-0339-900 with General Valve Iota One pulse driver) operating in 

parallel. A single spanwise row of holes was drilled into the suction 

surface of each blade at the inviscid pressure minimum location, 

s/Ls=0.5 (x/Cx=0.62), where s is the distance from the leading edge and 

Ls is the suction surface length. The pressure minimum has been 

shown in the studies listed above to be about the optimal location for 

flow control devices. The holes are 0.8 mm (0.006Cx) in diameter and 

tailboard

flap

flap

grid

B1

B2

B3

B4

B5

B6

B7

a) 

(b) 

Fig. 1 Drawings of test section: a) linear cascade, b) airfoil 

with VGJ holes and cross section of hole geometry 

drilled at 30° to the surface and 90° to the main flow direction, as 

shown in Fig. 1b. This is the same orientation used in all the VGJ 

studies listed above. The hole spacing is 10.6 diameters, and the 

length to diameter ratio is 12. With steady blowing and B=1, the mass 

flow rate of the jets is 0.04% of the main flow mass flow rate. With 

pulsed jets the mass flow is lower, particularly in low duty cycle cases.

 The solenoid valves pulse the VGJs, and the pulsing frequency is 

presented below in dimensionless form as F=fLj-te/Uave. For flow over 

single airfoils, F≥1 is typically needed to maintain separation control, 

but for cascades, Bons et al. [17] showed that control is possible in 

some cases with F=0.1. As shown in Volino [14] and Bons et al. [17], 

this is due to the extended calmed region which follows the jet 

disturbance. In practice, VGJs could be timed to wake passing in an 

LPT, which has a typical frequency of about F=0.3. 

 The center blade, designated B4 in Fig. 1, contains pressure taps 

near the spanwise centerline. Pressure surveys are made using a 

pressure transducer (0-870 Pa range Validyne transducer). Stagnation 

pressure is measured with a pitot tube upstream of the cascade. The 

uncertainty in the suction side pressure coefficients, Cp, is 0.07. Most 

of this uncertainty is due to bias error. Stochastic error is minimized 

by averaging pressure transducer readings over a 10 second period. 

 Total pressure losses are documented using a Kiel probe traversed 

across three blade spacings, 0.63 Cx downstream of the cascade. A 

traverse is located in the wind tunnel downstream of the cascade to 

move the probe. The traverse causes an acceptably low blockage 

when it is located at least two Cx downstream of the cascade. 

 Pressure and loss surveys were acquired at nominal Re=25,000 and 

50,000. The Reynolds number, as defined above, is based on the 

suction surface length and the nominal cascade exit velocity. The 

corresponding Reynolds numbers based on the cascade inlet velocity 

and the axial chord length are 10,000 and 20,000. 

NASA/CR—2012-217415 97
4 

 The jet velocity was measured at the exit plane of the VGJ holes and 

was documented in Volino et al. [29]. There is about a 5 ms delay 

between the opening of the solenoid valve and the beginning of the 

rise of the jet velocity above zero at the exit plane. The delay 

corresponds to the time required for a pressure wave to travel from the 

valve to the jet hole. The jet velocity rises rapidly and nearly 

continuously for another 6 ms to a maximum and then drops for about 

6 ms toward a steady value which is maintained for the rest of the time 

the valve is open. When the valve closes, the jet velocity drops to 

zero. In cases with high pulsing frequency or low duty cycle, the jet 

turns off before a steady velocity is reached. As will be shown below, 

the transient that occurs when the VGJ turns on is more effective for 

separation control than the steady blowing which may follow. The 

initial velocity maximum is therefore used to define the blowing ratio 

rather than the average velocity for the pulse. 

 Velocity profiles on the suction surface were measured at the six 

streamwise stations listed in Table 2. All stations are downstream of 

the inviscid pressure minimum at s/Ls=0.49. Profiles were acquired 

near the spanwise centerline of the airfoil with a hot-wire anemometer 

(AA Lab Systems model AN-1003) and a single sensor hot-film probe 

(TSI model 1201-20). The sensor diameter is 51 μm, and the active 

length is 1.02 mm. At each measurement location, data were acquired 

for 26 seconds at a 20 kHz sampling rate (219 samples). All raw data 

were saved. The high sampling rate provides an essentially continuous 

signal, and the long sampling time results in low uncertainty in both 

statistical and spectral quantities. Data were acquired at 40 wall 

normal locations in each profile, extending from the wall to the freestream, with most points concentrated in the near wall region. The 

probe was positioned as close to tangent to the airfoil surface as 

possible at each station, such that the probe body extended 

downstream of the sensor and the direction of the traverse was within 

5° of normal to the surface. In most cases the closest point to the wall 

in each profile was within about 0.2 mm of the wall, which compares 

to boundary layer thicknesses ranging from 1.1 mm to over 40 mm. 

 Flow direction in a separation bubble cannot be determined with a 

single-sensor hot-wire, but velocity magnitude can be measured and 

was found to be near zero within the bubbles of the present cases when 

the flow was laminar. In cases where the flow became turbulent but 

remained separated, fluctuating velocities caused false high mean 

velocity readings in the separation bubble. With the exception of these 

turbulent separated cases, the uncertainty in the mean velocity is 3-5% 

except in the very near wall region, where near-wall corrections (Wills 

[30]) were applied to the mean velocity. 

 Velocity data were both time averaged and ensemble averaged 

based on the phase within the jet pulsing cycle. Phase averages of 

mean and fluctuating velocity are shown below at 24 dimensionless 

times, t/T, within the pulsing cycle, where t is time and T is the period 

of the cycle. The time t/T=0 corresponds to the opening of the 

solenoid valve. 

 Wavelet spectra of the fluctuating velocity were computed using the 

method described in Volino [31]. In contrast to Fourier spectra, in 

which a signal is transformed from the time domain to the frequency 

domain, wavelet spectra provide the frequency content of a signal on a 

time resolved basis. The Mexican Hat wavelet was used for the 

analysis. The wavelet spectra were phase averaged to show the 

variation of frequency content in the velocity as a function of time 

within the jet pulsing cycle. 

NUMERICAL SIMULATIONS 

 Numerical simulations were conducted utilizing version 6.3.26 of 

the finite-volume code FluentTM[32]. The three dimensional 

computational domain includes a single passage. A uniform velocity 

Table 2: Velocity profile measurement stations 

Station 1 2 3 4 5 6 

s/Ls 0.53 0.59 0.69 0.78 0.88 0.97 

x/Cx 0.65 0.72 0.80 0.86 0.92 0.97 

inflow condition is specified 1.9Cx upstream of the blade leading edge 

in the flow direction. The inlet flow angle is set to 33o

 based on an 

inviscid calculation of the full cascade used in the experiment [27]. 

This angle agrees with the experimentally measured inlet angle to 

within the experiment uncertainty. The exit boundary is located 3.8Cx

downstream of the trailing edges in the flow direction. In the spanwise 

direction, the domain includes one VGJ. The boundary conditions on 

the sides of the passage are periodic. The full length of the hole is 

included in the simulations, allowing the jet velocity profile to develop 

before entering the main domain [29]. A uniform velocity boundary 

condition is specified at the hole inlet during jet blowing. 

Computation was continued until no variation cycle-to-cycle was 

reached. 

 Convergence was established when: 1) residuals reduced to a value 

10-5, 2) no change was observed in any field results, and 3) the mass 

imbalance was less than 0.01 %. 

 Large eddy simulations (LES) were done using the dynamic 

subgrid-scale kinetic energy model in Fluent, which is based on the 

model proposed by Kim and Menon [33]. In this model a separate 

transport equation is solved for subgrid-scale kinetic energy. The 

model constants are determined dynamically. The details of the 

implementation of this model in Fluent and its validation are given by 

Kim [34]. 

Code Validation 

 Three different grids were tested as shown in Table 3. Grid#37 was 

selected. Details of the grid structure are in Ibrahim et al. [35]. To 

accurately represent structures in the near-wall region (for LES) 

recommended values are: y

+

~2; Δx

+

~50-150; Δz

+

~15-40 (see Piomelli 

and Chasnov [36]). Grid#37 produced Cp versus s/Ls results in closest 

agreement to the experimental data and therefore was chosen for 

further computation. Computations with Grid#37 were then run for 

different time steps (0.0005, 0.0001 and 0.00005s) and time 

step=0.0001s was selected since not much improvement was achieved 

using the smaller step (0.00005s). 

RESULTS 

Re=25,000 

 Pressure profiles for cases with Re=25,000 and steady VGJ blowing 

are shown in Fig. 2. The inviscid profile for the L1A airfoil is shown 

as a common reference for comparing results between figures. At high 

Reynolds numbers (e.g. Re=300,000), experimental results agree with 

the inviscid line, as shown in Volino et al. [28]. The low peak 

followed by a plateau in the case without jets in Fig. 2 indicates 

separation without reattachment. As in the low freestream turbulence 

cases of Volino et al. [29] at this Reynolds number, blowing with 

B=2.0 or lower has no effect on separation. When B is increased to 2.5 

or 3.0, there are some signs of reattachment, but the Cp profile remains 

significantly different from the inviscid profile. 

 Figures 3 shows results for cases with pulsed jets and F=0.14. As in 

the low freestream turbulence cases, this frequency is too low to be 

very effective. The jets have some effect on Cp when B=3.0 and this 

effect increases at higher duty cycle. The total pressure losses for 

these cases are shown in Fig. 3b. A high Reynolds number 

(Re=300,000) baseline case from Volino et al. [28] is shown for 

comparison. The loss, ψ, is shown as a function of distance across the 

NASA/CR—2012-217415 98
5 

Table 3: Grid#3, 35 and 37 used in this investigation 

Grid # Size 

(Cells) 

Number of grids 

in z direction 

y+ Δz

+ Δx+

3 1,500,000 15 0.5 12.6 1–100 

35 5,900,000 30 0.5 6.3 0.4-52 

37 11,900,000 54 0.5 0.4-3.5 0.4-52 

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

Steady,B=1.0

Steady,B=2.0

Steady,B=2.5

Steady,B=3.0

Inviscid

Fig. 2 Cp profiles for steady blowing, Re=25,000 cases 

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

10%,B=2.0

10%,B=3.0

50%,B=1.0

50%,B=2.0

50%,B=3.0

Inviscid

a)

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.5

1

1.5

φ/Lφ

ψ

B3

B4

B5 B6

No control

10%,B=2.0

10%,B=2.5

10%,B=3.0

50%,B=2.0

50%,B=3.0

Steady,B=3.0

High Re

b) 

Fig. 3 Pressure results for F=0.14, Re=25,000 cases: a) Cp, 

b) total pressure loss 

cascade, φ, normalized on the blade spacing Lφ. The origin, φ=0, 

corresponds to the location downstream of the trailing edge of the 

center blade in the cascade (B4 in Fig. 1a) in the design flow direction. 

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

10%,B=1.0

10%,B=1.5

10%,B=2.0

10%,B=3.0

50%,B=1.0

50%,B=2.0

50%,B=2.5

50%,B=3.0

Inviscid

a)

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.5

1

1.5

φ/Lφ

ψ

B3

B4

B5 B6

No control

10%,B=1.5

10%,B=2.0

10%,B=3.0

50%,B=2.0

50%,B=3.0

High Re

b) 

Fig. 4 Pressure results for F=0.28, Re=25,000 cases: a) Cp, 

b) total pressure loss 

The pulsed jets reduce the loss significantly and shift the loss peaks to 

the right. This indicates that the separation bubble must be thinner, 

particularly when B=3.0, and that flow turning is increased. The loss 

peak at B=3.0 is narrower for the D=50% case than for the lower duty 

cycle, in agreement with the larger drop in Cp near the trailing edge 

for this case. Even in this best case, the loss peaks are significantly 

higher and to the left of the high Re case. 

 Results with F=0.28 are shown in Fig. 4. The jets begin to have an 

effect when B=1.5, and significant separation control is apparent when 

B=2.0 or higher, particularly with D=50%. The jets are not as 

effective when D=10%. The loss results show the VGJs at F=0.28 

have a strong effect. The duty cycle effect is less apparent than in the 

Cp results. 

 Figure 5 shows Cp results with F=0.56. The VGJs have some effect 

even at low blowing ratios, and when B=1.0 the boundary layer 

appears to reattach. The results do not appear to depend significantly 

on the duty cycle. This is consistent with the low freestream 

turbulence results, which showed that when F=0.56 the pulses occur 

frequently enough to control separation even at low duty cycles. 

Figure 6 shows results with F=1.12. Results are similar to those with 

F=0.56. Separation is controlled when B=0.75, and increasing B 

further does not appear to have much effect. Losses for the F=0.56 

and 1.12 cases are shown in Fig. 7. The losses are greatly reduced 

when separation is controlled, and with a few exceptions the loss 

results are consistent with the Cp results. 

 Integrated Cp and loss results are shown in Fig. 8. The Cp

difference between the pressure and suction sides of the passage is 

NASA/CR—2012-217415 99
6 

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

10%,B=0.5

10%,B=0.75

10%,B=1.0

10%,B=1.5

10%,B=2.0

10%,B=2.5

10%,B=3.0

50%,B=0.5

50%,B=1.0

50%,B=2.0

Inviscid

Fig. 5 Cp profiles for F=0.56, Re=25,000 cases 

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

10%,B=0.5

10%,B=0.75

10%,B=1.0

10%,B=1.5

10%,B=2.0

Inviscid

Fig. 6 Cp profiles for F=1.12, Re=25,000 cases 

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.5

1

1.5

φ/Lφ

ψ

B3

B4

B5 B6

No control

10%,B=0.75,F=0.56

10%,B=1.0,F=0.56

10%,B=2.0,F=0.56

50%,B=0.75,F=0.56

50%,B=1.0,F=0.56

50%,B=2.0,F=0.56

10%,B=1.0,F=1.12

10%,B=2.0,F=1.12

High Re

Fig. 7 Total pressure loss profiles for F=0.56 and F=1.12, 

Re=25,000 cases 

integrated in the axial direction to produce a quantity proportional to 

the lift on the airfoil,

∫

− =

Cx

0 x

suction pressure

int C

Cp( Cp dx) Cp (1) 

and this quantity is normalized using the same integrated quantity for 

the inviscid profile. The integrated total pressure loss, 

0 0.5 1 1.5 2 2.5 3

0.8

0.85

0.9

0.95

1

1.05

1.1

B

Cpint/Cpinv

Steady jets

F=0.14, D=10

F=0.14, D=50

F=0.28, D=10

F=0.28, D=50

F=0.56, D=10

F=0.56, D=50

F=1.12, D=10

a)

0 0.5 1 1.5 2 2.5 3

0

0.1

0.2

0.3

0.4

0.5

0.6

B

ψint

High Re

No jets

Steady jets

F=0.14 D=10

F=0.14 D=50

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

F=1.12 D=10

b)

0 0.5 1 1.5 2 2.5 3

0

2

4

6

8

10

12

14

16

B

Δθ [deg]

No jets

Steady jets

F=0.14 D=10

F=0.14 D=50

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

F=1.12 D=10

c) 

Fig. 8 Integrated pressure results for Re=25,000 cases: a) 

ratio of lift to lift in inviscid case, b) total pressure loss, c) 

change in exit flow angle from high Re case 

( )

( ) ST

Tej

1

j

2/L

2/L

int PP

PP

m

m

L

d

−

− = ∫ +

− 

φ 

φ φ

φψ ψ . (2) 

is an average of the loss profile result across one blade spacing 

centered on the peak corresponding to blade B4. As explained in 

Volino et al. [29], the loss associated with the jets themselves is 

included in the integrated loss. With steady jets the lift increases 

slowly as blowing ratio is increased, but never achieves the inviscid 

NASA/CR—2012-217415 100
7 

case value. The lift increases at each blowing ratio as F is increased, 

and is higher at the higher duty cycle. With F=0.56 and 1.12, the lift 

increases rapidly with B, reaching the invisid value when B=1, and 

then remains constant as B is increased further. In agreement with the 

lift, the integrated loss decreases with B from a high value of 0.55 

without flow control to about 0.22, with the best results at the higher 

pulsing frequencies. These values compare to a high Re value of about 

0.15. The losses with F=0.28 are somewhat lower than in the 

corresponding low TI cases of Volino et al. [29], but otherwise the 

high and low TI results are very similar. The change in exit flow angle 

is inferred from the shift in the location of the loss peaks from φ/Lφ=0 

in Figs. 3b, 4b and 7, and is shown as the difference from the high Re

case angle in Fig. 8c. The location of the peak in the high Re case 

corresponds to an exit flow angle of 54°, which agrees within the 

experimental uncertainty with the design exit flow angle of 57° 

corresponding to the design Zweifel=1.35. The design exit flow angle 

is lower than the 60° design metal angle of Table 1. Without flow 

control, turning is reduced by 16° compared to the high Re case. In 

the best controlled case, about 11° of turning is recovered, but there is 

still a 5.5° difference from the high Re case due to the much thicker 

boundary layer at low Re even with reattachment. 

 Time averaged velocity profiles for the baseline case and four cases 

with pulsed jets are shown in Fig. 9. The top row shows the mean 

velocity at the six streamwise stations noted in Table 2, and the lower 

row shows the rms fluctuating streamwise velocity, u′. Without flow 

control, the boundary layer has separated by Station 1 and the 

separation bubble grows at the downstream stations. The peak in u′ is 

in the shear layer far from the wall. With flow control it appears there 

may be a small separation bubble at the third and fourth stations, in 

agreement with the short plateau in Cp at this location in Figs. 4a and 

5, but it is much smaller than in the baseline case, and the boundary 

layer appears attached at the downstream stations. With F=0.28, 

increasing the blowing ratio from B=2 to 3 moves the u′ peak toward 

the wall, indicating a thinner bubble. The same trend is observed as B

in increased from 1 to 2 with F=0.56. Increasing F from 0.28 to 0.56 

results in significantly lower u′ in the outer part of the boundary layer, 

which is explained below. 

 Figure 10 shows phase averaged mean velocity profiles for the cases 

with F=0.28, D=10% and B=2.0, and with F=0.56, D=10% and B=2.0. 

The columns correspond to the six streamwise stations, and rows are 

for different phases in the pulsing cycle. A small separation bubble 

appears to form at Station 3 in the F=0.28 case, but is suppressed for 

several phases starting at t/T=0.417. The same is true at Station 4, 

with the suppression of the separation beginning at about t/T=0.5. At 

Station 5 the boundary layer appears to reattach at t/T=0.583, and 

remain attached until t/T=0.917. At station 6, reattachment occurs at 

about t/T=0.667, and the boundary layer remains attached until 

t/T=0.167 in the next cycle. 

 In the F=0.56 case there is less variation in the mean profile during 

the cycle. Variation in the velocity as the boundary layer separates 

and reattaches results in significant u′, and as the variation is reduced 

at higher F, u′ decreases. This is particularly apparent at y/Ls>0.1. As 

the separation bubble grows, fluid is forced away from the wall and 

the velocity at y/Ls>0.1 increases. When the boundary layer reattaches 

the velocity at y/Ls>0.1 drops. The result is high u′ in this region for 

the F=0.28 case in Fig. 9. The boundary layer appears to approach 

separation in the F=0.56 case for much of the cycle at Station 4, 

particularly between t/T=0.5 and 0.708, and at t/T=0.75 to 0.917 at 

Station 5 and t/T=0.083 at Station 6. At other times separation appears 

controlled. 

 Wavelet spectra help to explain the velocity profile results. Figure 

11 shows wavelet spectra for the F=0.28, D=10%, B=2.0 case. The 

0 1 0 1 0 1 0 1 0 1 0 1 0

0.1

0.2

U/Ue

y/Ls

St 12 3 4 5 6

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0

0.1

0.2

u′/Ue

y/Ls

No jets F=0.28 B=2.0 F=0.28 B=3.0 F=0.56 B=1.0 F=0.56 B=2.0

Fig. 9 Time averaged velocity profiles at six streamwise 

stations for Re=25,000 cases with no jets and jets with 

D=10%: top – mean velocity, bottom – rms velocity 

six plots in the figure correspond to the six streamwise stations. In 

each plot, the horizontal axis shows dimensionless frequency, fLs/Ue

on a log scale, and the vertical axis shows dimensionless time, t/T, for 

one pulsing cycle. Power spectral density is computed from 

instantaneous velocity data at all y locations and is shown in Fig. 11 

for the y locations corresponding to maximum time averaged u′ at each 

station. The contours show the power spectral density premultiplied 

by frequency and normalized by Ue

2

. The color scale is the same for 

all plots. The VGJ creates a disturbance at the beginning of a pulse, 

and the leading edge of this disturbance is visible as high contours 

centered at t/T=0.16, 0.25, 0.31, 0.38, 0.46 and 0.54 at Stations 1-6 

respectively. The arrival times at each station indicate that the leading 

edge of the disturbance convects along the surface at about 0.6 times 

the local freestream velocity. A second peak appears centered at 

t/T=0.24, 0.42, 0.54, 0.71, 0.93 and 0.06 at Stations 1-6 respectively. 

These peaks are believed to result when the trailing edge of the VGJ 

disturbance passes, and the times indicate a convection speed of about 

0.3 times the freestream velocity. Comparison of Figs. 10 and 11 

shows that at each station, the separation bubble is thickest at about the 

time the disturbance first arrives. Shortly after this, the velocity 

profile appears reattached. The profile remains attached at each 

station until about the time when the second peak appears in Fig. 11, 

indicating that the disturbance has past. The separation bubble then 

begins to grow almost immediately. Since the leading and trailing 

edges of the disturbance move at different velocities, the duration of 

the disturbance increases at the downstream stations. Hence, the 

duration of the separation bubble decreases at the downstream stations. 

If the pulsing frequency were increased, disturbances would follow 

each other more closely in time, and if the frequency were sufficiently 

high, the leading edge of each disturbance would overtake the trailing 

edge of the preceding disturbance, thereby eliminating the period of 

separated flow. Based on the convection speeds observed in Fig. 11, 

F=0.5 would be sufficient to eliminate separation at the trailing edge. 

 Low frequency peaks appear in Fig. 11 at the pulsing frequency, 

fLs/Ue=0.7, at all stations. A higher frequency peak at fLs/Ue=10 is 

also visible, particularly at Stations 1 and 2. The higher frequency is 

likely related to shear layer transition, and matches the frequency peak 

observed in transition without flow control in Volino et al. [28]. 

Figure 12 shows contours of the dimensionless, premultiplied wavelet 

spectra as a function of distance from the wall on the vertical axis and 

dimensionless frequency on the horizontal axis. The columns 

correspond to Stations 2 through 6, and the rows to different times 

during the cycle. The peak at fLs/Ue=0.7 appears at all times. Higher 

frequencies extending beyond fLs/Ue=10 are also present. The high 

frequency fluctuations have their highest magnitude in the shear layer 

NASA/CR—2012-217415 101
8 

0

0.1

0.083

t/T St. 1 2 3 4 5 6

0.083

0

0.1

0.167

0

0.1

0.25

0

0.1

0.333

0

0.1

0.417

0

0.1

0.5

y/Ls

0

0.1

0.583

0

0.1

0.667

0

0.1

0.75

0

0.1

0.833

0

0.1

0.917

0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

1

U/U

e

Fig. 10 Phase averaged mean velocity profiles for Re=25,000 cases, columns for six streamwise stations, rows for phases in 

pulsing cycle: blue – F=0.28, D=10%, B=2.0, red – F=0.56, D=10%, B=2.0 

NASA/CR—2012-217415 102
9 

t/T

St. 1

0.1 1 10 0

0.2

0.4

0.6

0.8

1

2

0.1 1 10

3

f⋅Ls

/Ue

0.1 1 10

4

0.1 1 10

5

0.1 1 10

6

0.1 1 10

Fig. 11 Wavelet spectra computed at y locations of maximum u′ in time averaged profiles and shown as function of time and 

frequency at six streamwise stations, F=0.28, D=10%, B=2.0, Re=25,000, solid white line is leading edge of disturbance, 

dashed white line is trailing edge 

St 2

0

0.2

0

0.2

0

0.2

0

0.2

0

0.2

y/Ls

0

0.2

0

0.2

0

0.2

0

0.2

0

0.2

0

0.2

0.1 1 10

0

0.2

3

0.1 1 10

4

f⋅Ls/Ue

0.1 1 10

5

0.1 1 10

6

0.083

t/T

0.167

0.25

0.333

0.417

0.5

0.583

0.667

0.75

0.833

0.917

1

0.1 1 10

Fig. 12 Wavelet spectra as function of distance from wall and frequency at Stations 2-6 (columns) and various phases (rows), 

F=0.28, D=10%, B=2.0, Re=25,000 

when the boundary layer is most separated and are suppressed when 

the VGJs cause reattachment. Another peak at fLs/Ue=2 and y/Ls

between 0.11 and 0.16 appears in the shear layer at t/T=0.417, 0.5 and 

0.583 at Stations 4 through 6 respectively. These are the arrival times 

of the VGJ disturbance at these stations. 

 Increasing the blowing ratio from 2 to 3 with F=0.28 increases the 

magnitude of the disturbances, but otherwise does not change the 

behavior observed in Figs. 9-12. The effect of pulsing frequency is 

shown in Figure 13 for the F=0.56, D=10%, B=2.0 case in the same 

format as Fig. 11. A peak frequency of about fLs/Ue=10 is still present 

at Stations 1 and 2, as in Fig. 11. High magnitudes are also visible at 

all stations at the pulsing frequency, which is now doubled to about 

fLs/Ue=1.4. In the format of Fig. 12 for this case (not shown), the 

arrival of the VGJ disturbance is again marked by a spectral peak in 

NASA/CR—2012-217415 103
10 

t/T

St. 1

0.1 1 10 0

0.2

0.4

0.6

0.8

1

2

0.1 1 10

3

f⋅Ls

/Ue

0.1 1 10

4

0.1 1 10

5

0.1 1 10

6

0.1 1 10

Fig. 13 Wavelet spectra computed at y locations of 

maximum u′ in time averaged profiles and shown as 

function of time and frequency at six streamwise stations, 

F=0.56, D=10%, B=2.0, Re=25,000, solid white line is leading 

edge of disturbance, dashed white line is trailing edge 

the shear layer at a frequency of about fLs/Ue=2. In Fig. 13, the center 

of the spectral peaks which presumably correspond to the first arrival 

of the VGJ disturbance appear at Stations 1-6 at t/T=0.17, 0.34, 0.5, 

0.67, 0.94 and 0.1 respectively. These dimensionless times are greater 

than those in Fig. 11 since the convection velocities are about the same 

in both cases but the pulsing period is halved. Comparing Figs. 13 and 

10, the separation bubble begins to grow at each station after the 

passing of the trailing edge of the disturbance, and the bubble is 

suppressed about 0.1T after the arrival of the next disturbance. This is 

the same behavior observed at F=0.28, but there is less time available 

for the bubble to grow at the higher frequency, so there is less 

variation of the velocity profile shape with time at F=0.56. The 

reduced separation at higher frequency is reflected in the lower losses 

and the Cp profiles, which are closer to the inviscid solution. 

Re=50,000 

 Pressure profiles for the cases with steady VGJs and Re=50,000 are 

shown in Fig. 14. The jets become effective when B≥1.5. Results for 

pulsed jets with F=0.14 are shown in Fig. 15. In terms of Cp, the jets 

begin to have an effect when B=0.75, but do not appear to fully control 

separation until B=1.5. Between B=0.75 and 1.5, a higher duty cycle 

appears to help, but at B=1.5 and above the results are about the same 

with D=10% or 50%. The total pressure loss profiles agree with the 

Cp results. In the best controlled case the loss peaks are still 

significantly higher and to the left of the high Re comparison case, but 

the difference is not as great as in the Re=25,000 cases above. Figure 

16 shows results with F=0.28. The jets have some effect with B=0.5 

and 0.75, and the effect increases at higher duty cycle. With B=1.0 

and higher, separation is controlled at both duty cycles. The loss peak 

is slightly narrower with D=50%. The Cp profiles show a small 

separation bubble still appears to remain at s/Ls=0.7. With F=0.56, as 

shown in Fig. 17, blowing with B=0.5 has some effect, particularly 

with D=50%, and with B=0.75 and D=10% the separation is 

controlled. Increasing B or D further has no additional effect. As in 

the lower frequency cases, a small separation bubble remains at 

s/Ls=0.7. The integrated Cp and loss results are shown in Fig. 18. The 

trends are the same as in the Re=25,000 cases of Fig. 8 and the low TI

cases of Volino et al. [29], but effective separation control and loss 

reduction is possible with somewhat lower B and F as Re or TI is 

increased. The reduction in flow turning compared to the high Re case 

is shown in Fig. 18c. Separation control increases the flow turning by 

about 12°, reducing the difference from the high Re case to 2°. 

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

Steady,B=0.5

Steady,B=0.75

Steady,B=1.0

Steady,B=1.5

Steady,B=2.0

Inviscid

Fig. 14 Cp profiles for steady blowing, Re=50,000 cases 

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

10%,B=0.25

10%,B=0.5

10%,B=0.75

10%,B=1.0

10%,B=1.5

10%,B=2.0

50%,B=0.25

50%,B=0.5

50%,B=0.75

50%,B=1.0

50%,B=1.5

50%,B=2.0

Inviscid

a)

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.5

1

1.5

φ/Lφ

ψ

B3

B4

B5 B6

No control

10%,B=1.0

10%,B=1.5

10%,B=2.0

50%,B=1.0

50%,B=1.5

High Re

b) 

Fig. 15 Pressure results for F=0.14, Re=50,000 cases: a) 

Cp, b) total pressure loss 

 Time averaged velocity profiles for some of the Re=50,000 cases 

are shown in Fig. 19. A large separation bubble without reattachment 

is present without flow control. Pulsed VGJs largely eliminate the 

bubble, although a small separation is still present at Station 3, in 

agreement with the small plateau in the Cp profiles of Figs. 16 and 17 

at s/Ls=0.7. Figure 20 shows phase averaged velocity profiles. With 

F=0.28, there is a small separation at t/T=0.417 at Station 4, at t/T=0.5 

at Station 5 and at t/T=0.583 at Station 6. Figure 21 shows the wavelet 

spectra for this case in the format of Fig. 11. The frequencies and 

times of the peaks in Figs. 21 and 11 are similar since the pulsing 

NASA/CR—2012-217415 104
11 

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

10%,B=0.25

10%,B=0.5

10%,B=0.75

10%,B=1.0

10%,B=1.5

50%,B=0.25

50%,B=0.5

50%,B=0.75

50%,B=1.0

Inviscid

a)

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.5

1

1.5

φ/Lφ

ψ

B3

B4

B5 B6

No control

10%,B=0.75

10%,B=1.0

50%,B=0.5

50%,B=0.75

50%,B=1.0

High Re

b) 

Fig. 16 Pressure results for F=0.28, Re=50,000 cases: a) 

Cp, b) total pressure loss 

frequency, F=0.28, is the same for both cases. As in the lower Re

case, the boundary layer reattaches shortly after the appearance of the 

VGJ disturbance at each station. In contrast to the lower Re case, the 

boundary layer remains attached for about 0.2T after the disturbance 

has passed, resulting in attached flow for most of the pulsing cycle. 

The F=0.56, D=10%, B=0.75 case in Fig. 20 shows little variation of 

the velocity profiles during the pulsing cycle. The spectra for this 

case, shown in Fig. 22 show nearly a continuous disturbance through 

the cycle. There may be some signs of separation with a thin bubble at 

Station 3, but at the downstream stations the near overlap of the 

disturbance period between consecutive pulses combined with the 

lesser tendency toward separation at Re=50,000 than at 25,000, results 

in an attached boundary layer for the full cycle. 

Computational Results 

 Figure 23 shows Cp for the Re=50,000, B=1, D=10%, F=0.28 case. 

The spike at the suction peak is at the jet location. In agreement with 

the experimental data, the CFD shows a small separation bubble 

occurs but the flow remains reattached over most of the airfoil. Figure 

24 show time averaged mean velocity at six streamwise stations 

comparing the CFD results with the experimental data. The agreement 

is good upstream, but differences are apparent at the downstream 

stations, particularly near the wall. The CFD predicts an exit flow 

angle of 57° both in the present case and in a high Re case. The exit 

angle difference between the high Re and controlled Re=50,000 case 

was also small in the experimental result in Fig. 18c. The 57° agrees 

with the design exit flow angle, but is a few degrees higher than the 

0 0.2 0.4 0.6 0.8 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

10%,B=0.25

10%,B=0.5

10%,B=0.75

10%,B=1.0

50%,B=0.25

50%,B=0.5

50%,B=0.75

Inviscid

a)

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.5

1

1.5

φ/Lφ

ψ

B3

B4

B5 B6

No control

10%,B=0.5

10%,B=0.75

10%,B=1.0

50%,B=0.5

50%,B=0.75

High Re

b) 

Fig. 17 Pressure results for F=0.56, Re=50,000 cases: a) 

Cp, b) total pressure loss 

experimental value. Figure 25 shows the Q-Criterion, which is used to 

illustrate vortices [32], colored by axial velocity, Vx, at different times 

in the cycle. At the beginning of the blowing there is no separation 

bubble present near the trailing edge of the airfoil. This is in contrast 

with the corresponding low TI case (see Ibrahim et al. [35]). However 

there is a small bubble located midway between the jet and the trailing 

edge. During blowing that separated region travels downstream while 

the flow is attached along most of the airfoil. Right after shutdown of 

the jet the separation region is smaller. Also at that time the shear 

layer is highly energized. Near the end of the cycle the flow is 

attached at the trailing edge, but a small separated region begins to 

appear again upstream. These results are consistent with the 

experimental velocity profiles of Fig. 20, which show a small 

separation bubble moving down the airfoil between wakes, but 

attached flow at all locations for most of the cycle. 

DISCUSSION 

 The VGJ disturbances convect downstream in the shear layer over 

the separation bubble, and if they are of sufficient amplitude, cause the 

boundary layer to attach at a given streamwise location shortly after 

the arrival of the disturbance. The boundary layer tends to remain 

attached until after the passage of the trailing edge of the disturbance, 

although at the lowest Reynolds numbers there may be some tendency, 

as shown in Fig. 10, to separate earlier. After the disturbance passes, 

the separation bubble begins to re-grow. This happens almost 

immediately at very low Reynolds number and more slowly at higher 

Re. The trailing edge of each disturbance travels slower than the 

NASA/CR—2012-217415 105
12 

0 0.5 1 1.5 2

0.8

0.85

0.9

0.95

1

1.05

B

Cpint/Cpinv

Steady jets

F=0.14, D=10

F=0.14, D=50

F=0.28, D=10

F=0.28, D=50

F=0.56, D=10

F=0.56, D=50

a)

0 0.5 1 1.5 2

0

0.1

0.2

0.3

0.4

0.5

0.6

B

ψint

High Re

No jets

F=0.14 D=10

F=0.14 D=50

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

b)

0 0.5 1 1.5 2 0

2

4

6

8

10

12

14

16

B

Δθ [deg]

No jets

F=0.14 D=10

F=0.14 D=50

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

c) 

Fig. 18 Integrated pressure results for Re=50,000 cases: a) 

ratio of lift to lift in inviscid case, b) total pressure loss, c) 

change in exit flow angle from high Re case 

leading edge, resulting in a stretching of the disturbance in the 

streamwise direction as it moves downstream and a longer duration of 

the disturbance at downstream locations. The result is a shorter period 

between disturbances at downstream stations and less separated flow. 

Increasing the pulsing frequency reduces the period between 

disturbances at all locations and results in better separation control. 

 The present results, in agreement with all the previous VGJ studies 

referenced above, show that pulsed jets are more effective than steady 

jets. The disturbance created when a jet is first turned on is more 

effective than any steady blowing which may follow. Velocity results 

from cases with low freestream turbulence (not shown in the present 

0 1 0 1 0 1 0 1 0 1 0 1 0

0.1

0.2

U/Ue

y/Ls

St 12 3 4 5 6

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0

0.1

0.2

u′/Ue

y/Ls

No jets F=0.28 B=1.0 F=0.56 B=0.75

Fig. 19 Time averaged velocity profiles at six streamwise 

stations for Re=50,000 cases with no jets and jets with 

D=10%: top – mean velocity, bottom – rms velocity 

paper) show that when pulsing with D=50%, the boundary layer may 

re-separate during the period of steady blowing. Increasing the duty 

cycle from 10 to 50% can, however, provide better separation control, 

particularly at low pulsing frequencies, as shown above and in Volino 

et al. [29]. Velocity profiles and spectra from the low TI, D=50% 

cases show a second effective disturbance occurs each cycle when the 

VGJs turns off, thereby doubling the disturbance frequency above the 

pulsing frequency. In the D=10% cases, the disturbances created 

when the jets turn on and off occur so close together in time that they 

effectively constitute a single disturbance. 

 The present results are very similar to the low TI results of Volino et 

al. [29]. High TI does help to control separation, allowing VGJ flow 

control with slightly lower amplitude and frequency jets in some cases, 

but the physics of the separation and flow control appears to be the 

same at high and low TI. Increasing the VGJ velocity and frequency 

help to reduce the separation bubble extent both temporally and 

spatially, and this increases lift and reduces total pressure losses. Once 

the separation is largely eliminated, increasing the blowing ratio or 

pulsing frequency further provides no additional benefit. 

CONCLUSIONS 

 The effect of vortex generator jets on the flow over the very high lift 

L1A airfoil was studied under high freestream turbulence conditions. 

Reynolds numbers based on suction surface length and nominal exit 

velocity of 25,000 and 50,000 were considered. Without flow control, 

the boundary layer separated and did not reattach. Flow control with 

VGJs was possible even at Re=25,000. In agreement with previous 

studies, pulsed jets were more effective than steady jets. Pulsing with 

a dimensionless frequency of F=0.56 or higher allowed for separation 

control with blowing ratio of at most 1.0 and 10% duty cycle. 

Effective separation control resulted in a 20% increase in lift and up to 

a 60% reduction in total pressure loss compared to baseline cases at 

the same Reynolds number. Loss values still remain higher than in 

high Reynolds number cases. Pulsed jets at lower frequencies 

provided partial separation control and some loss reduction and could 

be improved somewhat by increasing the blowing ratio or duty cycle. 

 Phase averaged velocity profiles and wavelet spectra show the 

boundary layer intermittently reattaching as disturbances pass and then 

separating between disturbances. Increasing the pulsing frequency 

reduces the time available for separation. When the time available is 

sufficiently small, the boundary layer remains attached at all times. At 

Re=25,000, separation was nearly fully controlled for the full pulsing 

cycle when F=0.5. Higher frequency pulsing provided little additional 

benefit. At higher Reynolds numbers the separation bubble grows 

more slowly, so lower pulsing frequencies can be tolerated. 

NASA/CR—2012-217415 106
13 

0

0.1

0.083

t/T St. 1 2 3 4 5 6

0.083

0

0.1

0.167

0

0.1

0.25

0

0.1

0.333

0

0.1

0.417

0

0.1

0.5

y/Ls

0

0.1

0.583

0

0.1

0.667

0

0.1

0.75

0

0.1

0.833

0

0.1

0.917

0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

1

U/U

e

Fig. 20 Phase averaged mean velocity profiles for Re=50,000 cases, columns for six streamwise stations, rows for phases in 

pulsing cycle: blue – F=0.28, D=10%, B=1.0, red – F=0.56, D=10%, B=0.75 

NASA/CR—2012-217415 107
14 

t/T

St. 1

0.1 1 10 0

0.2

0.4

0.6

0.8

1

2

0.1 1 10

3

f⋅Ls

/Ue

0.1 1 10

4

0.1 1 10

5

0.1 1 10

6

0.1 1 10

Fig. 21 Wavelet spectra computed at y locations of 

maximum u′ in time averaged profiles and shown as 

function of time and frequency at six streamwise stations, 

F=0.28, D=10%, B=1.0, Re=50,000, solid white line is leading 

edge of disturbance, dashed white line is trailing edge t/T

St. 1

0.1 1 10 0

0.2

0.4

0.6

0.8

1

2

0.1 1 10

3

f⋅Ls

/Ue

0.1 1 10

4

0.1 1 10

5

0.1 1 10

6

0.1 1 10

Fig. 22 Wavelet spectra computed at y locations of 

maximum u′ in time averaged profiles and shown as 

function of time and frequency at six streamwise stations, 

F=0.56, D=10%, B=0.75, Re=50,000, solid white line is 

leading edge of disturbance, dashed line is trailing edge 

0.00

0.50

1.00

1.50

2.00

2.50

0.00 0.10 0.20 0.30 0.40 0.50 0.60 0.70 0.80 0.90 1.00

s/Ls

Cp

CFD

Experiment

Fig. 23 Cp profile for B=1, F=0.28, D=10%, Re=50,000 

s/Ls = 0.53

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

-0.5 0 0.5 1 1.5

U/Ue

y/Ls

CFD

Exp.

s/Ls = 0.59

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

-0.5 0 0.5 1 1.5

U/Ue

y/Ls

CFD

Exp.

s/Ls = 0.69

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

-0.5 0 0.5 1 1.5

U/Ue

y/Ls

CFD

Exp.

s/Ls = 0.78

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

-0.5 0 0.5 1 1.5

U/Ue

y/Ls

CFD

Exp.

s/Ls = 0.88

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

-0.5 0 0.5 1 1.5

U/Ue

y/Ls

CFD

Exp.

s/Ls = 0.97

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

-0.5 0 0.5 1 1.5

U/Ue

y/Ls

CFD

Exp.

Fig 24 U/Ue at for B=1, F=0.28, D=10%, Re=50,000 

 Computational results from a large eddy simulation show 

reattachment in agreement with the experiments for a case with 

effective VGJ flow control. The LES results allow visualization of 

structures such as vortices, providing insight into the flow behavior. 

 High freestream turbulence makes the boundary layer less likely to 

separate and allows for separation control with slightly lower blowing 

ratios and pulsing frequencies than in low TI cases. The differences 

observed between high and low TI were small, however, and the 

physics of the separation and flow control appear to be the same in 

both cases. 

ACKNOWLEDGMENTS

 This work was sponsored by the National Aeronautics and Space 

Administration under grant NNC07IA10I. The grant monitors were 

Drs. Anthony Strazisar and James Heidmann of the NASA Glenn 

Research Center. The support of the United States Naval Academy 

Technical Support Department Shop and Fluids Laboratory is greatly 

appreciated. We greatly appreciate the computer time provided for us 

by the Ohio Super Computer (OSC). The OSC Computer Cluster has 

been made available as part of the Center’s mission to support Ohio 

Universities. 

NASA/CR—2012-217415 108
15 

Fig 25 Q-Criterion contours colored by axial velocity, B=1, 

F=0.28, D=10%, Re=50,000, at different times in the cycle 

 

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[10] Bons, J.P., Hansen, L.C., Clark, J.P., Koch, P.J., and 

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[13] Volino, R.J., 2003, “Passive Flow Control on Low-Pressure 

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[14] Volino, R.J., 2003, “Separation Control on Low-Pressure Turbine 

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of Turbomachinery, 125, pp. 765-777. 

[15] Huang, J., Corke, T., and Thomas, F., 2003, “Plasma Actuators 

for Separation Control on Low Pressure Turbine Blades,” AIAA 

Paper 2003-1027. 

[16] Johnston, J.P., and Nishi, M., 1990, “Vortex Generator Jets. 

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[17] Bons, J.P., Sondergaard, R., and Rivir, R.B., 2002, “The Fluid 

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NASA/CR—2012-217415 109
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[19] McQuilling, M., and Jacob, J., 2004, “Effect of Chord Location 

on Separation Control With Vortex Generator Jets on Low 

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Melwani, N., 1997, “Investigation of the Calmed Region Behind 

a Turbulent Spot,” ASME Journal of Turbomachinery, 119, pp. 

802-809. 

[22] Schulte, V., and Hodson, H.P., 1998, “Prediction of the Becalmed 

Region for LP Turbine Profile Design,” ASME Journal of 

Turbomachinery, 120, pp. 839-846. 

[23] Bons, J.P., Reimann, D., and Bloxham, M., 2008, “Separated 

Flow Transition on an LP Turbine Blade With Pulsed Flow 

Control,” ASME Journal of Turbomachinery, 130, 021014. 

[24] Clark J.P., 2007, Private Communication, Air Force Research 

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[25] Zhang, X.F., and Hodson, H., 2005, “Combined Effects of 

Surface Trips and Unsteady Wakes on the Boundary Layer 

Development of an Ultra-High-Lift LP Turbine Blade,” ASME 

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2008, “The Application of Flow Control to an Aft-Loaded Low 

Pressure Turbine Cascade with Unsteady Wakes,” ASME Paper 

GT2008-50864.

[27] Ibrahim, M., Kartuzova, O., and Volino, R.J., 2008, 

“Experimental and Computational Investigations of Separation 

and Transition on a Highly Loaded Low Pressure Turbine Airfoil: 

Part 1 – Low Freestream Turbulence Intensity,” ASME Paper 

IMECE2008-68879. 

[28] Volino, R.J, Kartuzova, O., and Ibrahim, M., 2008, 

“Experimental and Computational Investigations of Separation 

and Transition on a Highly Loaded Low Pressure Turbine Airfoil: 

Part 2 – High Freestream Turbulence Intensity,” ASME Paper 

IMECE2008-68776.

[29] Volino, R.J., Kartuzova, O., and Ibrahim, M.B., 2009, 

“Experimental and Computational Investigations of Low-Pressure 

Turbine Separation Control Using Vortex Generator Jets,” ASME 

Paper GT2009-59983. 

[30] Wills, J.A.B., 1962, “The Correction of Hot-Wire Readings for 

Proximity to a Solid Boundary,” Journal of Fluid Mechanics, 12, 

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[31] Volino, R.J., 2005, “An Investigation of the Scales in Transitional 

Boundary Layers under High Freestream Turbulence 

Conditions,” Experiments in Fluids, 38, pp. 516-633. 

[32] Fluent, Inc., 2005, Fluent 6.3 - User Guide. 

[33] Kim, W.W., and Menon, S., 1997, “Application of the localized 

dynamic subgrid-scale model to turbulent wall-bounded flows,” 

AIAA Paper AIAA-97-0210. 

[34] Kim, S.E., 2004, “Large eddy simulation using unstructured 

meshes and dynamic subgridscale turbulence models,” AIAA 

Paper AIAA-2004-2548. 

[35] Ibrahim, M.B., Kartuzova, O., Doucet, D., and Volino, R.J., 

2010, “LES Flow Control Simulations for Highly Loaded Low 

Pressure Turbine Airfoil (L1A) Using Pulsed Vortex Generator 

Jets,” ASME Paper GT2010-23015. 

[36] Piomelli, U., and Chasnov, J.R., 1995, “Large-Eddy Simulations: 

Theory and Applications,” Turbulence and Transition Modeling, 

lecture notes from the ERCOFTAC/IUTAM summer school, 

Stockholm, 12-20 June, 1995. 

NASA/CR—2012-217415 110
1 

Proceedings of ASME Turbo Expo 2010: Power for Land, Sea and Air 

GT2010 

June 14-18, 2010, Glasgow, UK

GT2010-23573 

EFFECT OF UNSTEADY WAKES ON BOUNDARY LAYER SEPARATION ON A VERY 

HIGH LIFT LOW PRESSURE TURBINE AIRFOIL 

Ralph J. Volino

Mechanical Engineering Department 

United States Naval Academy 

Annapolis, Maryland 21402-5042 

volino@usna.edu 

ABSTRACT 

 Boundary layer separation has been studied on a very high lift, lowpressure turbine airfoil in the presence of unsteady wakes. 

Experiments were done under low (0.6%) and high (4%) freestream 

turbulence conditions on a linear cascade in a low speed wind tunnel. 

Wakes were produced from moving rods upstream of the cascade. 

Flow coefficients were varied from 0.35 to 1.4 and wake spacing was 

varied from 1 to 2 blade spacings, resulting in dimensionless wake 

passing frequencies F=fLj-te/Uave (f is the frequency, Lj-te is the length 

of the adverse pressure gradient region on the suction surface of the 

airfoils, and Uave is the average freestream velocity) ranging from 0.14 

to 0.56. Pressure surveys on the airfoil surface and downstream total 

pressure loss surveys were documented. Instantaneous velocity profile 

measurements were acquired in the suction surface boundary layer and 

downstream of the cascade. Cases were considered at Reynolds 

numbers (based on the suction surface length and the nominal exit 

velocity from the cascade) of 25,000 and 50,000. In cases without 

wakes, the boundary layer separated and did not reattach. With wakes, 

separation was largely suppressed, particularly if the wake passing 

frequency was sufficiently high. At lower frequencies the boundary 

layer separated between wakes. Background freestream turbulence 

had some effect on separation, but its role was secondary to the wake 

effect. 

NOMENCLATURE 

Cp ( ) 2

T − ρU/PP2 e , pressure coefficient 

Cx axial chord length 

F fLj-te/Uave, dimensionless frequency 

f frequency 

Lj-te length of adverse pressure gradient region on suction surface

Ls suction surface length 

Lφ blade spacing (pitch) 

P pressure 

PS upstream static pressure 

PT upstream stagnation pressure 

PTe downstream stagnation pressure 

Re UeLs/ν, exit Reynolds number 

s streamwise coordinate, distance from leading edge 

T period of jet pulsing cycle 

t time 

TI background freestream turbulence intensity 

U local mean velocity 

Uave average freestream velocity in adverse pressure gradient region 

Ui inlet freestream velocity 

Ue nominal exit freestream velocity, based on inviscid solution 

u′ rms fluctuating streamwise velocity 

x axial distance from leading edge 

φ coordinate along blade spacing, normal to axial chord 

ν kinematic viscosity 

ρ density 

ψ (PT-PTe)/(PT-PS), total pressure loss coefficient 

ζ Uicos(αi)/Urod = Uaxial/Urod, flow coefficient

INTRODUCTION 

 Partial loss of lift and higher aerodynamic losses can be caused by 

boundary layer separation on low-pressure turbine (LPT) airfoils (e.g. 

Hourmouziadis [1], Mayle [2], and Sharma et al. [3]). In aircraft 

engines the lower density and therefore lower Reynolds numbers at 

altitude can lead to a component efficiency drop of 2% between 

takeoff and cruise in large commercial transport engines, and possibly 

as much as 7% in smaller engines operating at higher altitudes [4, 5]. 

Separation can occur when airfoil loading is too high because of the 

strong adverse pressure gradients on the suction surface. High loading 

is, however, desirable since it can be used to reduce airfoil count, 

weight and cost. The challenge is to design high lift airfoils which do 

not have separation problems. This requires accurate prediction of 

suction side separation under relevant conditions. 

 Separation can be strongly affected by wakes shed from the airfoils 

in upstream stages in an engine. The velocity deficit and elevated 

turbulence in periodic wakes help to suppress separation and can cause 

a separated boundary layer to reattach. Hodson and Howell [6] 

describe the mechanisms by which wakes promote reattachment, 

including the “negative jet” which results when the velocity deficit in 

the wake causes the flow outside the wake to accelerate and impinge 

on the surface, and the unsteadiness which promotes transition in the 

boundary layer. Following the wake itself is a calmed period 

(Gostelow et al. [7] and Schulte and Hodson [8]) in which the 

NASA/CR—2012-217415 111

This material is declared a work of the U.S. Government and is not subject to copyright protection in the United States. Approved for public release; 

distribution is unlimited. 
2 

boundary layer has low turbulence but is resistant to separation. 

Numerous studies have considered the wake effect in the LPT, 

including those listed in Hodson and Howell [6], and more recent 

references in Bons et al. [9] and Pluim et al. [10]. Examples include 

Schobeiri et al. [11], Öztürk and Schobeiri [12], Jiang and Simon [13], 

and Mahallati and Sjolander [14] who all used the Pack B airfoil. 

Zhang and Hodson [15] and Funazaki et al. [16] used more highly 

loaded airfoils. Many additional studies are available from these 

research groups and others. 

 Airfoils can be designed with high resistance to separation, as 

described by Praisner and Clark [17], and knowledge of wake effects 

allows for designs with higher loading than would be possible under 

steady inflow conditions. Even with wakes, however, a loading limit 

will always exist, above which separation will still occur. 

 In the present study, a very highly loaded airfoil that exhibits 

separation even in the presence of wakes is used. The airfoil, known 

as the L1A, was designed at the Air Force Research Laboratory 

(AFRL) and is available on a limited basis from Clark [18]. 

Dimensions of the L1A as used in the present study are given in Table 

1. Based on the design calculations of Clark [18], the L1A has a 

Zweifel coefficient of 1.35, which corresponds to 10% higher loading 

than the “ultra-high lift” airfoils described by Zhang and Hodson [19], 

and 17% higher loading than the Pack B airfoil. The L1A is aft 

loaded, which is advantageous for reducing secondary flow losses at 

the endwalls, but makes the boundary layer more prone to separation 

than a forward loaded blade, as documented in Bons et al. [9], Volino 

[20], Ibrahim et al. [21], and Volino et al. [22]. In cases without 

wakes and low Reynolds numbers, the boundary layer separates and 

does not reattach, in spite of transition to turbulence in the shear layer 

over the separation bubble. This result contrasts with the results of 

studies on less aggressive airfoils (e.g. Volino [23]), which all showed 

reattachment after transition. The separation bubble on the L1A is 

about four times thicker than that on the Pack B. The larger distance 

from the shear layer to the wall on the L1A apparently prevents the 

turbulent mixing in the shear layer from reaching the wall and causing 

reattachment. The failure of the boundary layer to reattach results in a 

20% loss in lift and increases profile losses by up to a factor of 7. At 

higher Reynolds numbers the separation bubble closes, and for 

Re>200,000 the separation bubble is small and the boundary layer is 

attached over most of the airfoil. 

 The effect of wakes on the L1A boundary layer was studied by 

Bons et al. [9], who considered a case with Re=50,000 (based on the 

suction surface length and the nominal exit velocity from the cascade), 

background freestream turbulence TI=3%, and periodic wakes 

produced with moving rods upstream of the airfoils. The 

dimensionless frequency of the wake passing was F=fLj-te/Uave=0.34, 

where Lj-te is the length of the adverse pressure gradient region on the 

suction surface, and Uave is the average freestream velocity over this 

distance. The length Lj-te is also the distance from a row of vortex 

generator jet (VGJ) holes to the trailing edge. Pulsed vortex generator 

jets were used by Bons et al. [9] and Volino et al. [24] to control 

separation on the L1A. Although these jets are not considered in the 

present work, the same dimensionless frequency, F, is used to describe 

the wake passing in the present study and the jet pulsing in [9] and 

[24]. Bons et al. [9] found that wakes reduced the separation bubble 

thickness significantly and reduced total pressure losses by more than 

50%, but did not cause boundary layer reattachment. This result 

contrasts with earlier work on less aggressive airfoils (e.g. Reimann et 

al. [25] on the Pack B) which showed reattachment for at least part of 

the wake passing period, or a reduction in bubble size in cases with 

already reattached boundary layers (e.g. Schobeiri et al. [11]). The 

much thicker separation bubble on the L1A prevents the mixing 

Table 1: Cascade parameters 

Axial 

Chord, Cx

[mm] 

True 

Chord 

[mm] 

Pitch, 

Lφ

[mm] 

Span 

[mm] 

Suction 

side, Ls

[mm] 

Inlet 

flow 

angle 

Exit 

flow 

angle 

134 146 136 724 203 35° 60° 

associated with the wake from penetrating all the way to the airfoil 

surface and causing full reattachment. 

 In the present study, the effect of wakes is considered with 

Re=25,000 and 50,000 under low (0.6%) and high (4%) freestream 

turbulence conditions. Cases with various wake passing frequencies 

and flow coefficients are documented. Surface pressure distributions, 

total pressure loss profiles, and instantaneous boundary layer velocity 

measurements are used to explain the mechanism by which the wakes 

affect separation. 

EXPERIMENTAL FACILITY AND MEASUREMENTS

 Experiments were conducted in a closed loop wind tunnel with a 

seven blade linear cascade as shown in Fig. 1. A fine screen located 

upstream of the cascade is used to break up the boundary layers which 

form upstream of the test section and to provide uniform inlet 

conditions to the cascade. The freestream turbulence entering the 

cascade was measured with a cross-wire probe positioned just 

upstream of the center blade. The turbulence intensity is 0.8% in the 

streamwise component and 0.5% in the cross stream components. The 

integral length scale of the streamwise component is 6.3 cm. To 

produce high freestream turbulence, the screen is replaced with a 

coarse grid, consisting of a 1.5 mm thick sheet metal plate with 19 mm 

square holes spaced 25.4 mm apart, center to center, in both directions. 

In a plane perpendicular to the inlet flow and 1.7Cx upstream of the 

center blade, the grid produced uniform flow with TI=6.0% in the 

streamwise component and 4.2% in the cross stream components, for 

an overall intensity of 4.9%. The streamwise component was also 

measured at the inlet plane of the cascade in the four center passages, 

where it had decayed to between 4 and 4.2% between blades B2 and 

B5 and 4.6% between blades B5 and B6. Downstream of the cascade, 

the local TI is 1.8% across all passages, as documented in Volino et al. 

[22]. The local freestream turbulence intensity in the passage at the 

beginning of the adverse pressure gradient region is 1.4%. The change 

in TI through the passage is due mainly to the change in the local 

freestream velocity along with some decay of the turbulence. The 

upstream integral length scale of the freestream turbulence is 0.12Cx in 

the streamwise component and 0.04Cx in the other components. 

Further details of the facility and inlet flow are in Volino et al. [22]. 

 A tailboard, shown in Fig. 1, is needed to produce the correct exit 

flow angle from the cascade. Its position was set to produce 

periodicity at high Reynolds numbers. A tailboard on the opposite 

side of the cascade and inlet guide vanes were found to be 

unnecessary. To produce the correct approach flow to the end blades 

(B1 and B7), the amount of flow escaping around the two ends of the 

cascade is controlled with the flaps shown in Fig. 1. The inlet flow 

angle was checked with a three-hole pressure probe and found to agree 

with the design angle to within 2° of uncertainty. Good periodicity at 

high Reynolds numbers was obtained in the exit flow. At low 

Reynolds numbers, when significant separation bubbles are present, 

the periodicity is not as good due to suppression of the separation 

bubble thickness on the blades closest to the tailboard. In cases where 

wakes or other flow control suppress separation, periodicity is 

reestablished. The lack of periodicity in cases with large separation 

bubbles is considered acceptable since the focus of the research is 

separation control, and not documentation of cases with large 

separation that would be unacceptable in practice. This compromise 

NASA/CR—2012-217415 112
3 

tailboard

flap

flap

grid

B1

B2

B3

B4

B5

B6

B7

Fig. 1 Schematic of linear cascade 

facilitates the study of a larger number of cases by obviating the need 

to adjust the tailboard by trial and error for each case. It also provides 

for better repeatability in the experiments, since the position of the 

tailboard is fixed for all cases. Any changes in separation with wakes 

will be larger in practice than documented in the experiment, due to 

the effect of the tailboard in suppressing the bubble size in the nowake cases. 

 The wake generator includes a chain near each endwall of the 

cascade that passes 0.54Cx upstream of the leading edges of the 

cascade blades. The chains then pass downstream around blade B7 on 

the inside turn of the cascade and pass well downstream of the cascade 

before returning upstream around blade B1 on the outside turn of the 

cascade. This completes the chain circuit. A traverse for probe 

movement is located within the chain circuit downstream of the blade 

row. Each chain is driven by a drive gear and also passes around six 

idler sprockets. One of the idler sprockets is adjustable to maintain 

tension in the chain. The drive gears for the upper and lower chains 

are on a common axle and driven by a single electric motor so both 

chains move in unison. The motor speed is set with a variable 

frequency inverter. The chain links have hollow pins, through which 

the wake generator rods are attached. Each rod consists of a 4 mm 

diameter carbon fiber tube with a steel pin attached at each end. The 

steel pins are inserted through the holes in the upper and lower chain, 

and then secured with small clips. Tests were run with average 

distances between rods of 136 mm, 221 mm, and 272 mm, which 

correspond to 1Lφ, 1.6Lφ and 2Lφ, where Lφ is the blade spacing in the 

cascade. These ratios of rod to blade spacing are in the range expected 

for vane to rotor blade spacing in an engine. 

 The ratio of the rod diameter to the axial chord is 0.03, which is 

consistent with the wake generators of Bons et al. [9] and Funazaki et 

al. [16]. The rods are smaller than those of Kaszeta et al. [26] who had 

a diameter to chord ratio of 0.06. The present rods are larger than 

those of Schobeiri et al. [11] and Zhang and Hodson [15] who had rod 

diameter to chord ratios of about 0.01. In the present study, as in Bons 

et al. [9] and Kaszeta et al. [26], the rod wakes are intended to simulate 

the wakes of very highly loaded airfoils under low Reynolds number 

conditions with thick boundary layers and in some cases large 

separation bubbles. A large diameter rod is therefore needed to 

simulate an airfoil wake with a large velocity deficit. The studies 

using smaller diameter rods were done at higher Reynolds numbers, so 

separation bubbles were smaller and rods producing lower velocity 

deficits were more appropriate. 

 For most tests, the rods were driven at a velocity of 1.18 times the 

cascade inlet velocity, Ui. This gives a flow coefficient, 

ζ =Uicos(αi)/Urod=0.7, where αi is the inlet flow angle. This is in the 

expected range for an engine. The timing of the wake generator is 

recorded with an infrared photo detector, which senses the passage of 

each rod and emits a voltage that is recorded simultaneously as other 

data are acquired, allowing phase averaging of the results. 

Measurements 

 The center blade, designated B4 in Fig. 1, contains pressure taps 

near the spanwise centerline. Pressure surveys are made using a 

pressure transducer (0-870 Pa range Validyne transducer). Stagnation 

pressure is measured with a pitot tube upstream of the cascade and 

wake generator. The uncertainty in the suction side pressure 

coefficients, Cp, is 0.07. Most of this uncertainty is due to bias error. 

Stochastic error is minimized by averaging pressure transducer 

readings over a 10 second period. 

 Total pressure losses are documented using a Kiel probe traversed 

across three blade spacings, 0.63Cx downstream of the cascade. A 

traverse is located in the wind tunnel downstream of the cascade to 

move the probe. The traverse causes an acceptably low blockage 

when it is located at least two Cx downstream of the cascade. 

 Pressure and loss surveys were acquired at nominal Re=25,000 and 

50,000. The Reynolds number, as defined above, is based on the 

suction surface length and the nominal cascade exit velocity. The 

corresponding Reynolds numbers based on the cascade inlet velocity 

and the axial chord length are 10,000 and 20,000. 

 Velocity profiles on the suction surface were measured at the six 

streamwise stations listed in Table 2. All stations are downstream of 

the inviscid pressure minimum at s/Ls=0.49. Profiles were acquired 

near the spanwise centerline of the airfoil with a hot-wire anemometer 

(AA Lab Systems model AN-1003) and a single sensor hot-film probe 

(TSI model 1201-20). The sensor diameter is 51 μm, and the active 

length is 1.02 mm. At each measurement location, data were acquired 

for 26 seconds at a 20 kHz sampling rate (219 samples). All raw data 

were saved. The high sampling rate provides an essentially continuous 

signal, and the long sampling time results in low uncertainty in both 

statistical and spectral quantities. Data were acquired at 40 wall 

normal locations in each profile, extending from the wall to the freestream, with most points concentrated in the near wall region. The 

probe was positioned as close to tangent to the airfoil surface as 

possible at each station, such that the probe body extended 

downstream of the sensor and the direction of the traverse was within 

5° of normal to the surface. In most cases the closest point to the wall 

in each profile was within about 0.2 mm of the wall, which compares 

to boundary layer thicknesses ranging from 1.1 mm to over 40 mm. 

 Flow direction in a separation bubble cannot be determined with a 

single-sensor hot-wire, but velocity magnitude can be measured and 

was found to be near zero within the bubbles of the present cases when 

the flow was laminar. In cases where the flow became turbulent but 

remained separated, fluctuating velocities caused false high mean 

velocity readings in the separation bubble. With the exception of these 

turbulent separated cases, the uncertainty in the mean velocity is 3-5% 

except in the very near wall region, where near-wall corrections (Wills 

[27]) were applied to the mean velocity. 

 Velocity was also measured downstream of the cascade along the 

same line used for the total pressure loss measurements. Downstream 

and boundary layer velocity data were both time averaged and 

ensemble averaged based on the phase within the wake passing period. 

Phase averages of mean and fluctuating velocity are shown below at 

24 dimensionless times, t/T, within the wake passing period, where t is 

time and T is the period between wakes. 

NASA/CR—2012-217415 113
4 

Table 2: Velocity profile measurement stations 

Station 1 2 3 4 5 6 

s/Ls 0.53 0.59 0.69 0.78 0.88 0.97 

x/Cx 0.65 0.72 0.80 0.86 0.92 0.97 

RESULTS 

Wake Characteristics 

 Several studies have compared rod wakes to airfoil wakes. 

Schobeiri et al. [28] discussed the theory for describing wake flows. 

Pfeil and Herbst [29] showed that the far wake of an airfoil has almost 

the same wake as a cylinder having the same drag, although the rods 

have no means of matching the correct potential field of adjacent blade 

rows [30, 31]. Pluim et al. [10] tested rods of various shapes and 

provide references to earlier studies. They found that the wakes from 

wedge shaped rods agreed best with L1A airfoil wakes, but cylindrical 

rods also provided a reasonable approximation. Figure 2 shows the 

mean and rms fluctuating streamwise velocity in the wakes of the 

moving rods and the cascade blades of the present study. The cascade 

wakes were measured by traversing a hot-wire probe in lines 0.3Cx and 

0.63Cx downstream of the trailing edges of the blades, perpendicular to 

the axial direction. The variable φ denotes the distance across the 

cascade. The origin, φ=0, is directly downstream of the center blade 

(B4) in the design flow direction, while φ/Lφ = -1 and 1 are 

downstream of blades B3 and B5 respectively. The spacing between 

airfoil rows in an engine is expected to be about 0.3Cx according to 

Pluim et al. [10], so this location would be where the wake from an 

upstream airfoil would reach the leading edge of the next row. The 

moving rod wakes were measured by placing a stationary hot-wire 

midway between blades B4 and B5 in the plane of the leading edges. 

The moving rods traversed 0.54Cx upstream of the hot-wire. The data 

from the probe were phase averaged on the rod passing cycle to 

determine the mean and fluctuating velocity as functions of time. For 

comparison in the figures, the rod velocity was then used to convert 

time to distance traversed. Curves in the figures were shifted along the 

horizontal axis to align the peaks for comparison. All curves were 

normalized using the average local mean velocity. 

 The wakes from the upstream rods, when shown in dimensionless 

form as in Fig. 2, did not depend on Reynolds number. For all of the 

flow and rod velocities considered, the Reynolds number based on rod 

diameter is between about 400 and 2000, which should result in 

laminar separation for the boundary layer on the rod and similar wakes 

for all cases. The wake for low TI, Re=50,000, ζ=0.7, 1.6Lφ rod 

spacing, and dimensionless wake passing frequency of F=0.28 is 

shown. In contrast to the rod wakes, the wakes of the cascade airfoils 

depend strongly on Reynolds number, as shown in Volino [20]. Due 

to a large separation bubble, the airfoil wake at Re=50,000 has a large 

mean velocity deficit at 0.3Cx which decreases 80% in peak magnitude 

by 0.63Cx. The peak turbulence intensity drops by 40% over this 

distance. At Re=200,000 the boundary layer remains attached, so the 

airfoil wake is about half as wide as at the lower Re, and the peak 

velocity deficit is about 25% as large. The peak turbulence intensity is 

7%, which is about a third that at the lower Re. The change in the 

wake between 0.3Cx and 0.63Cx is much smaller than at the lower Re. 

The rod wake lies between the high and low Re airfoil wakes. Its peak 

turbulence intensity is 14%. The velocity deficit and turbulence level 

in the rod wakes in Fig. 2 appear to be reasonable approximations for 

an airfoil wake, although they do not fully match either the high or low 

Re case. One could conceivably use a rod of either larger or smaller 

diameter to match the low or high Re airfoil cases better. In cases with 

high freestream turbulence, the rod wakes are nearly the same as those 

in Fig. 2, with 14% peak turbulence intensity. The only difference is 

−0.5 −0.4 −0.3 −0.2 −0.1 0 0.1 0.2 0.3 0.4 0.5 0.5

0.6

0.7

0.8

0.9

1

1.1

1.2

1.3

φ/Lφ

U/U

mean

Rod wake

Cascade wake, Re=50,000, 0.3Cx

Cascade wake, Re=50,000, 0.63Cx

Cascade wake, Re=200,000, 0.3Cx

Cascade wake, Re=200,000, 0.63Cx

a)

−0.5 −0.4 −0.3 −0.2 −0.1 0 0.1 0.2 0.3 0.4 0.5 0

0.05

0.1

0.15

0.2

0.25

φ/Lφ

u′/U

mean

Rod wake

Cascade wake, Re=50,000, 0.3Cx

Cascade wake, Re=50,000, 0.63Cx

Cascade wake, Re=200,000, 0.3Cx

Cascade wake, Re=200,000, 0.63Cx

b)

Fig. 2 Velocity in wakes of wake generator rods and 

cascade wakes with low TI: a) mean, b) rms 

between the wakes, where the background turbulence rises to 4% with 

high TI. 

 The cases of the present study will be used for comparison to future 

cases with wakes and vortex generator jet flow control. Figure 3 

compares the upstream rod wake to the airfoil wake in a case with 

Re=50,000, and pulsed VGJ blowing with F=0.56, 10% duty cycle and 

blowing ratio (jet velocity/freestream velocity) equal 1. The VGJs 

provided good flow control in this case. The boundary layer separates 

very briefly but reattaches. Further details are available in Volino et 

al. [24]. For the airfoil wake in Fig. 3, the phase averaged results at 24 

phases through the VGJ pulsing cycle are shown. Since the boundary 

layer approaches separation during part of the cycle, the wake velocity 

profile magnitude and position changes somewhat during the cycle. 

The mean velocity deficit and wake width of the upstream rod and 

airfoil are about the same. The peak turbulence intensity in the rod 

wake is about 40% higher than the highest value in the airfoil wake. 

 Turbulence spectra provide additional information about the wakes. 

Spectra of the fluctuating streamwise velocity were computed for the 

cascade airfoil wakes and the rod wakes. Without flow control, the 

airfoil wakes at Re=50,000 exhibit sharp, high amplitude, low 

frequency (F=0.7) peaks associated with the shedding frequency of the 

separation bubble. With Re=200,000, the amplitude of the peaks is 

about two orders of magnitude lower than at the lower Re, and the 

peaks are broad banded and centered at F=2.5. These peaks are likely 

associated with the turbulence in the airfoil boundary layers. The 

spectra of the upstream rod wakes have amplitude and frequency 

between the high and low Re cases, agreeing most closely with the 

Re=50,000 VGJ controlled case mentioned above. Figure 4 shows 

NASA/CR—2012-217415 114
5 

−0.5 −0.4 −0.3 −0.2 −0.1 0 0.1 0.2 0.3 0.4 0.5 0.7

0.75

0.8

0.85

0.9

0.95

1

1.05

1.1

1.15

φ/Lφ

U/U

mean

Rod wake

Cascade wake

a)

−0.5 −0.4 −0.3 −0.2 −0.1 0 0.1 0.2 0.3 0.4 0.5 0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

0.2

φ/Lφ

u′/U

mean

Rod wake

Cascade wake

b)

Fig. 3 Velocity in wakes of wake generator rods and phase 

averaged velocity in wakes of cascade blades for low TI, 

Re=50,000 case with VGJ flow control [24]: a) mean, b) rms 

wavelet spectra (computed as described in Volino [32] using the 

Mexican Hat wavelet) for the airfoil and rod wakes. Power spectral 

density premultiplied by frequency is shown as a function of 

dimensionless frequency and distance across the cascade. One 

representative phase within the VGJ pulsing cycle is shown for the 

airfoil wake. As in Figs. 2 and 3, time is converted to position for the 

rod wake. For the airfoil wake there is a large peak at F=1.8 with 

magnitude 0.0052 and a smaller peak at F=0.6 with magnitude 0.0029 

that is hidden behind the larger peak in Fig. 4a. The magnitude of the 

two peaks varies during the VGJ cycle. The rod wake has a large peak 

at F=0.8 with magnitude 0.0047 and a smaller peak at F=0.2 with 

magnitude 0.0016. Figs. 2-4 show that the velocity deficit, turbulence 

intensity and spectral content are of the same order for the airfoil and 

rod wakes, although the agreement with any one case is not perfect. 

Since the wake of an upstream airfoil will not necessarily match 

exactly the wake of the airfoil on which it impinges, the present rod 

wakes are considered to provide a reasonable approximation to the 

wakes that might be shed from an upstream blade row at low Reynolds 

numbers. 

Low TI 

Re=25,000. The Cp and total pressure loss profiles for cases with 

Re=25,000 and low TI are shown in Fig. 5. The inviscid Cp profile for 

the L1A airfoil is shown for comparison. The low peak followed by a 

plateau in the case without wakes indicates separation without 

reattachment. For the cases with ζ=0.7, decreasing the rod spacing 

from 2Lφ to 1Lφ (which corresponds to increasing the wake passing 

−1.5

−1

−0.5

0

0.5

1

1.5

10−2

10−1

100

101

102

0

1

2

3

4

5

6

7

x 10−3

φ/Lφ

F

f⋅PSD/U2

e

a)

0

0.2

0.4

0.6

0.8

1

1.2

1.4

10−2

10−1

100

101

102

0

1

2

3

4

5

x 10−3

φ/Lφ

F

f⋅PSD/U2

e

b)

Fig. 4: Wavelet spectra of wake velocity: a) cascade blades 

for low TI, Re=50,000 case with VGJ flow control [24], b) 

wake generator rods 

frequency, F from 0.22 to 0.45) causes Cp to approach the inviscid 

line and reduces the loss by over 40% compared to the no wake case. 

The wakes also cause the loss peaks to shift to the right in Fig. 4b, 

indicating greater flow turning resulting from a thinner separation 

bubble. The greater turning agrees with a 20% increase in lift 

indicated by the Cp profiles. Also shown in Fig. 5 is a case with 1.6Lφ

rod spacing and the rod speed and F doubled so ζ=0.35. The higher 

wake passing frequency helps to reduce the separation. 

 Figures 6-9 show time-space plots of the phase averaged mean and 

fluctuating velocity 0.63Cx downstream of the blade row for the cases 

of Fig. 5. The contours are normalized by the exit velocity Ue. In Fig. 

6, the rod spacing is 2Lφ, so the flow in alternating passages is in 

phase, with the passages between a half cycle out of phase. The 

turbulence (Fig. 6b) shows vertical strips of moderately high 

fluctuations at φ/Lφ = -1.5, -0.5, 0.5 and 1.5 associated with the airfoil 

wakes. These positions correspond to the loss peaks in Fig. 5b. 

Between these peaks the fluctuation levels range from very low 

(corresponding to the low background TI) to high when the rod wakes 

pass. The rod wakes proceed at an angle in the figure, rising from left 

to right, as they move forward in time and transit across the cascade. 

The highest turbulence peaks occur where the rod wakes interact with 

the separation bubble and airfoil wakes. At the position and time of 

each rod wake peak in the turbulence is a corresponding velocity 

deficit in the mean velocity (Fig. 6a). Following the velocity deficit is 

a mean velocity maximum as the flow accelerates after the wake 

passes. In Fig. 7, the rod spacing is reduced to 1.6Lφ. All passages are 

out of phase with each other, and the peaks are stretched along the 

NASA/CR—2012-217415 115
6 

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No wakes

2Lφ, ζ=0.7

1.6Lφ, ζ=0.7

1Lφ, ζ=0.7

1.6Lφ, ζ=0.35

Inviscid

a)

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4

B5 B6

No wakes

2Lφ, ζ=0.7

1.6Lφ, ζ=0.7

1Lφ, ζ=0.7

1.6Lφ, ζ=0.35

b)

Fig. 5 Pressure profiles for low TI, Re=25,000 cases: a) Cp, 

b) Total pressure loss 

time axis compared to Fig. 6 since the rod speed is the same but the 

wake passing period is shorter. In Fig. 8, the rod and blade spacing are 

equal, so all passages are in phase with each other. The wake passing 

frequency is high enough that there is little time for the boundary layer 

to separate between wakes. This reduces the peak turbulence level. 

There is still a mean velocity deficit associated with each rod wake 

followed by a velocity rise, but the difference between the minima and 

maxima is reduced since suppression of the separation bubble reduces 

the flow unsteadiness. Figure 9 shows the effect of increasing the rod 

speed. Again, the increased wake passing frequency helps to suppress 

the separation and reduce unsteadiness. 

 Figure 10 shows time averaged mean and fluctuation velocity 

profiles at the six streamwise stations of Table 2. Without wakes the 

separation bubble is very thick, and the peak fluctuating velocity is in 

the shear layer over the bubble. As shown in Volino [20], the shear 

layer does transition to turbulent between Stations 3 and 4, but this 

does not cause reattachment. Wakes with rod spacing of 2Lφ and 1.6Lφ

reduce the separation bubble thickness by about half, but do not 

eliminate it. Figure 11 shows phase averaged mean velocity for the 

2Lφ spacing case. The six columns correspond to the six streamwise 

stations, and the rows correspond to different phases in the wake 

passing cycle. The boundary layer separates with a very thin bubble at 

Station 2 that becomes thicker at the downstream stations at most 

phases. The bubble appears as a region of nearly constant, but nonzero velocity near the wall. The non-zero velocity is an artifact of the 

hot-wire sensor’s inability to distinguish flow direction and the 

turbulent fluctuations within the bubble. There is also a period of 

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

a)

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

b)

Fig. 6 Time space plot of phase averaged velocity 0.63Cx

downstream of cascade for low TI, Re=25,000, ζ=0.7, 2Lφ

rod spacing: a) U/Ue, b) u′/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

a)

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

b)

Fig. 7 Time space plot of phase averaged velocity 0.63Cx

downstream of cascade for low TI, Re=25,000, ζ=0.7, 1.6Lφ

rod spacing: a) U/Ue, b) u′/Ue

NASA/CR—2012-217415 116
7 

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

a)

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

b) 

Fig. 8 Time space plot of phase averaged velocity 0.63Cx

downstream of cascade for low TI, Re=25,000, ζ=0.7, 1Lφ

rod spacing: a) U/Ue, b) u′/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

a)

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

b)

Fig. 9 Time space plot of phase averaged velocity 0.63Cx

downstream of cascade for low TI, Re=25,000, ζ=0.35, 1.6Lφ

rod spacing: a) U/Ue, b) u′/Ue

0 1 0 1 0 1 0 1 0 1 0 1 0

0.1

0.2

U/Ue

y/Ls

St 12 3 4 5 6

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0

0.1

0.2

u′/Ue

y/Ls

No wakes 2Lφ 1.6Lφ

Fig. 10 Time averaged velocity profiles at six streamwise 

stations for low TI, Re=25,000, ζ=0.7 cases: top – mean, 

bottom – rms 

0

0.1

0.083

t/T St. 1 2 3 4 5 6

0

0.1

0.167

0

0.1

0.25

0

0.1

0.333

0

0.1

0.417

0

0.1

0.5

y/Ls

0

0.1

0.583

0

0.1

0.667

0

0.1

0.75

0

0.1

0.833

0

0.1

0.917

0 1 0 1 0 1 0 1 0 1 0 1 0

0.1

1

U/Ue

Fig. 11 Phase averaged mean velocity profiles for low TI, 

Re=25,000, ζ=0.7, 2Lφ rod spacing; columns for six 

streamwise stations, rows for phases in wake cycle 

attached flow at each station where the velocity far from the wall is 

somewhat lower than at other phases and the near wall velocity more 

smoothly approaches zero, as fluid from the outer flow fills in the 

region near the wall when the separation bubble collapses. The 

reattachment occurs at t/T=0.125-0.25, t/T=0.167-0.333, t/T=0.25-

0.458, t/T=0.375-0.583, and t/T=0.458-0.708 at Stations 2-6 

respectively. 

 More results for the low TI, Re=25,000 cases are shown in Fig. 12 

as contour plots of the phase averaged velocity at different phases 

within the wake passing cycle. The vertical axis shows distance from 

NASA/CR—2012-217415 117
8 

y/Ls 0.05

0.1

0.15

0.2

y/Ls

t/T=0.042

0.05

0.1

0.15

0.2

t/T=0.125 t/T=0.208 t/T=0.292

y/Ls 0.05

0.1

0.15

0.2

y/Ls

s/Ls

t/T=0.375

0.6 0.8

0.05

0.1

0.15

0.2

s/Ls

t/T=0.625

0.6 0.8 s/Ls

t/T=0.875

0.6 0.8 s/Ls

t/T=0.958

0.6 0.8

a)

y/Ls 0.05

0.1

0.15

0.2

y/Ls

t/T=0.042

0.05

0.1

0.15

0.2

t/T=0.208 t/T=0.292 t/T=0.375

y/Ls 0.05

0.1

0.15

0.2

y/Ls

s/Ls

t/T=0.458

0.6 0.8

0.05

0.1

0.15

0.2

s/Ls

t/T=0.542

0.6 0.8 s/Ls

t/T=0.625

0.6 0.8 s/Ls

t/T=0.833

0.6 0.8

b)

U/Ue

0 0.5 1 1.5

u′/Ue

0 0.1 0.2 0.3

Fig. 12 Phase averaged mean (rows 1 and 3) and rms 

streamwise fluctuating (rows 2 and 4) velocity for low TI, 

Re=25,000, ζ=0.7: a) 2Lφ rod spacing, b) 1.6Lφ rod spacing 

the wall and the horizontal axis shows distance along the suction 

surface. The first and third row in each case show mean velocity, and 

the second and fourth row show the corresponding fluctuating 

velocity. Figure 12a shows the same case as Fig. 11. The effect of the 

rod wake appears in the mean velocity as elevated freestream velocity 

at t/T=0.958. The location of high freestream velocity is farther 

downstream at t/T=0.042, and reaches the trailing edge at about 

t/T=0.292. During this time the boundary layer remains thick, as 

shown by the blue color in the figure, but the near wall velocity is 

slightly elevated (lighter blue). The higher near wall velocity 

corresponds to the attached profiles at these phases in Fig. 11. As the 

wake passes, the fluctuating velocity rises in the freestream and near 

the wall. Between t/T=0.375 and 0.625 the boundary layer relaxes to 

an undisturbed, separated state, and then continues in this state until 

the next wake arrives. Figure 12b shows results with the wake spacing 

reduced to 1.6Lφ. The wake passing is similar to that in Fig. 12a, but 

the dimensionless duration of the wake passing is longer since the 

wake travels at the same velocity but the period is shorter. This results 

in less time for recovery between wakes. 

 Figure 13 presents another view of the wake passing effect as timespace plots of momentum thickness. The horizontal axis shows 

distance along the suction surface and the vertical axis shows 

dimensionless time. The data are repeated for two cycles to show the 

periodicity in time. The white lines in the figure show the 

approximate location of the wake and the calmed region following the 

wake. The leading edge of the wake affected region in the boundary 

s/Ls

t/T

0.55 0.6 0.65 0.7 0.75 0.8 0.85 0.9 0.95

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

1

2

3

4

5

6

7

8

9

a) 

s/Ls

t/T

0.6 0.65 0.7 0.75 0.8 0.85 0.9 0.95

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

0

1

2

3

4

5

6

7

b) 

Fig. 13 Phase averaged momentum thickness for low TI, 

Re=25,000, ζ=0.7: a) 2Lφ rod spacing, b) 1.6Lφ rod spacing; 

lines show leading edge (solid) and trailing edge (dash-dot) 

of wake, and trailing edge of calm region (dashed) 

layer travels at about the freestream velocity, while the trailing edges 

of the wake and calmed region move at about 50% and 30% of the 

freestream velocity respectively. These convection velocities agree 

with values in the literature, such as those reported by Stieger and 

Hodson [33] and Zhang and Hodson [15]. Momentum thickness is 

shown since it is proportional to losses in the boundary layer. Plots of 

other quantities such as the shape factor also show the wake and 

calmed regions, as shown in Stieger and Hodson [33]. In Fig. 13a, the 

wake causes a strip of elevated momentum thickness corresponding to 

the reattachment and elevated near wall turbulence in Figs. 11 and 12a. 

The periods of high momentum thickness and turbulence produce the 

high turbulence peaks observed in the airfoil wake in Fig. 6. In Fig. 

13, during the early part of the calmed period the momentum thickness 

is slightly suppressed compared to the value between wake events. 

During the calmed period the disturbance has already passed, as shown 

in Fig. 12a, but the boundary layer remains attached, as shown in Fig. 

11. Figure 13b shows results with 1.6Lφ wake spacing. As in Fig. 

13a, the momentum thickness increases as the wake passes and is 

reduced in the calm region. With less time for recovery between 

wakes, the variation in momentum thickness is lower in Fig. 13b. 

 The present results are similar to those in other studies with wakes, 

such as Schobeiri et al. [11] and Zhang and Hodson [15]. In those 

studies the boundary layer reattached even without wakes, and the 

passing wake caused a reduction in the separation bubble size. In the 

present study, because the Reynolds numbers are lower and the 

pressure gradients stronger, the boundary layer only intermittently 

reattaches as the wake passes. There are several mechanism by which 

wakes could induce reattachment. The elevated freestream turbulence 

in the wake will cause transition to move upstream in the shear layer, 

and if transition occurs before the separation bubble becomes too 

NASA/CR—2012-217415 118
9 

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No wakes

2Lφ, ζ=0.7

1.6Lφ, ζ=0.7

1Lφ, ζ=0.7

1.6Lφ, ζ=1.4

Inviscid

a)

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4

B5 B6

No wakes

2Lφ, ζ=0.7

1.6Lφ, ζ=0.7

1Lφ, ζ=0.7

1.6Lφ, ζ=1.4

b)

Fig. 14 Pressure profiles for low TI, Re=50,000 cases: a) 

Cp, b) Total pressure loss 

thick, the turbulent mixing could extend close enough to the wall to 

cause reattachment. Volino et al. [22] showed, however, that elevated 

background turbulence alone did not cause reattachment on the L1A at 

Re=25,000. The upstream turbulence intensity in the wake is 14%, as 

documented above, but as the flow is accelerated through the passage, 

this local intensity drops to about 3% at the beginning of the adverse 

pressure gradient region. Separated flow transition correlations, such 

as those of Roberts and Yaras [34] and Volino and Bohl [35], indicate 

that a rise to 3% TI would cause transition to shift upstream by about 

0.06Ls, which may not be enough to explain the full wake effect on 

reattachment. Other effects of wakes are explained in detail by Stieger 

and Hodson [33] and Zhang and Hodson [15], who used instantaneous 

velocity field measurements to document negative jets resulting from 

temporal freestream acceleration and rollup vortices in the shear layer 

which caused reattachment. These are likely the same mechanisms 

responsible for reattachment in the present study. 

Re=50,000. Pressure results for low TI cases with Re=50,000 are 

shown in Fig. 14. Without wakes the shear layer does not reattach, as 

in the Re=25,000 case. With wakes and ζ=0.7, reattachment is clear 

for all rod spacings and losses are reduced by over 40%. With the rod 

speed reduced so ζ=1.4 (F reduced to 0.14), only partial reattachment 

is observed. Wake velocity is shown in Figs. 15-17 for the ζ=0.7 

cases with rod spacings of 2Lφ, 1.6Lφ and 1Lφ respectively. The airfoil 

wakes, appearing as vertical strips of low mean velocity and about 7% 

turbulence, are in the expected positions based on the loss peaks of 

Fig. 14b. There is some rise in the magnitude of the turbulence to 

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

a)

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

b)

Fig. 15 Time space plot of phase averaged velocity 0.63Cx

downstream of cascade for low TI, Re=50,000, ζ=0.7, 2Lφ

rod spacing: a) U/Ue, b) u′/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

a)

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

b)

Fig. 16 Time space plot of phase averaged velocity 0.63Cx

downstream of cascade for low TI, Re=50,000, ζ=0.7, 1.6Lφ

rod spacing: a) U/Ue, b) u′/Ue

NASA/CR—2012-217415 119
10 

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

a)

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

b)

Fig. 17 Time space plot of phase averaged velocity 0.63Cx

downstream of cascade for low TI, Re=50,000, ζ=0.7, 1Lφ

rod spacing: a) U/Ue, b) u′/Ue

about 10% when the airfoil and rod wakes combine in each cycle, 

particularly in the cases with farther rod spacing. Between the airfoil 

wakes, the rod wakes appear in the freestream once per cycle with 

about 5% turbulence. The mean velocity drops during the rod wakes 

and then rises after the wake passes. The variability in the velocity 

and turbulence are much lower than in the Re=25,000 cases because 

the wakes suppress the separation bubble more at the higher Re. This 

is shown in the time averaged velocity profiles in Fig. 18. Without 

wakes there is a large separation bubble. With wakes the boundary 

layer is attached. The phase averaged mean velocity profiles are 

shown in Fig. 19 for the ζ=0.7, 2Lφ (F=0.22) case. The wakes affect 

the freestream velocity, particularly at the upstream stations, and a thin 

separation bubble is visible at Stations 3 and 4 between wakes. At the 

downstream stations the boundary layer appears attached at all times. 

Figure 20a shows contours of the phase averaged mean and fluctuating 

velocity for the ζ=0.7, 2Lφ case. The velocity variation in the 

freestream is visible between t/T=0.042 and 0.375 as the wake passes, 

and the freestream turbulence is slightly elevated during this period as 

well. The boundary layer turbulence also rises as the wake passes. 

The variation during the cycle is not as great as in the Re=25,000 cases 

of Fig. 12. The boundary layer remains relatively thin and attached at 

all phases. The time-space plot of the momentum thickness is shown 

for this case in Fig. 21a. As in the lower Re cases, the wake causes an 

increase in momentum thickness as it passes. Any delay in 

momentum thickness growth in the calmed region is slight. The 

overall value of the momentum thickness is less than half the value in 

the Re=25,000 cases, which corresponds to the lower losses in Fig. 14 

compared to Fig. 5 for the ζ=0.7, 2Lφ cases and the lower wake 

turbulence in Fig. 15 compared to Fig. 6. A case was also documented 

with ζ=0.7 and 1.6Lφ. It showed little difference from the ζ=0.7, 2Lφ 

0 1 0 1 0 1 0 1 0 1 0 1 0

0.1

0.2

U/Ue

y/Ls

St 12 3 4 5 6

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0

0.1

0.2

u′/Ue

y/Ls

No wakes 2Lφ 1.6Lφ

Fig. 18 Time averaged velocity profiles at six streamwise 

stations for low TI, Re=50,000, ζ=0.7 cases: top – mean, 

bottom – rms 

0

0.1

0.083

t/T St. 1 2 3 4 5 6

0

0.1

0.167

0

0.1

0.25

0

0.1

0.333

0

0.1

0.417

0

0.1

0.5

y/Ls

0

0.1

0.583

0

0.1

0.667

0

0.1

0.75

0

0.1

0.833

0

0.1

0.917

0 1 0 1 0 1 0 1 0 1 0 1 0

0.1

1

U/Ue

Fig. 19 Phase averaged mean velocity profiles for low TI, 

Re=50,000, ζ=0.7 cases, 2Lφ rod spacing; columns for six 

streamwise stations, rows for phases in wake cycle 

case. A reduction of the wake rod velocity to ζ=1.4, 1.6Lφ (F=0.14) 

produces the results in Figs. 20b and 21b. With the reduced wake 

frequency the boundary layer is separated for much of the cycle and 

the momentum thickness is over double its value in the ζ=0.7 case. 

These results agree with the high loss and separated flow Cp profile 

with ζ=1.4 in Fig. 14. 

High TI 

 The pressure results for the high freestream turbulence cases with 

Re=25,000 are shown in Fig. 22. Wake velocity contours are shown in 

NASA/CR—2012-217415 120
11 

y/Ls 0.05

0.1

0.15

0.2

y/Ls

t/T=0.042

0.05

0.1

0.15

0.2

t/T=0.125 t/T=0.208 t/T=0.292

y/Ls 0.05

0.1

0.15

0.2

y/Ls

s/Ls

t/T=0.375

0.6 0.8

0.05

0.1

0.15

0.2

s/Ls

t/T=0.625

0.6 0.8 s/Ls

t/T=0.875

0.6 0.8 s/Ls

t/T=0.958

0.6 0.8

a)

y/Ls 0.05

0.1

0.15

0.2

y/Ls

t/T=0.042

0.05

0.1

0.15

0.2

t/T=0.208 t/T=0.375 t/T=0.542

y/Ls 0.05

0.1

0.15

0.2

y/Ls

s/Ls

t/T=0.667

0.6 0.8

0.05

0.1

0.15

0.2

s/Ls

t/T=0.833

0.6 0.8 s/Ls

t/T=0.917

0.6 0.8 s/Ls

t/T=1

0.6 0.8

b)

U/Ue

0 0.5 1 1.5

u′/Ue

0 0.1 0.2 0.3

Fig. 20 Phase averaged mean (rows 1 and 3) and rms 

streamwise fluctuating (rows 2 and 4) velocity for low TI, 

Re=50,000: a) ζ=0.7, 2Lφ, b) ζ=1.4,1.6Lφ

Figs. 23 and 24. The results are similar to the corresponding low TI

results of Figs. 5, 7 and 8. The Cp distributions indicate slightly less 

separation with high TI, and in agreement the mean velocities show 

less variation and the turbulence values are slightly lower with high TI. 

 Results for the Re=50,000 cases are shown in Figs. 25-27. For the 

ζ=0.7 cases, the results are nearly the same as in the corresponding 

low TI cases of Figs. 14, 16 and 17. In the 1.6Lφ, ζ=1.4 cases 

separation and losses are higher than with lower ζ, but are clearly 

reduced by the high TI. The effect of the rod wakes appears to 

dominate over the background freestream turbulence effect, and if the 

wake passing frequency is high enough to suppress separation, the TI

level between wakes is largely irrelevant. If there is sufficient time 

between wakes for the boundary layer to separate, high background TI

helps to limit the separation. 

 In agreement with results from the literature for other airfoils, wakes 

help to suppress boundary layer separation. For the L1A airfoil with 

Re=25,000, a dimensionless wake passing frequency of about F=0.5 

appears sufficient to effectively eliminate separation. At lower 

frequencies the boundary layer separates between wakes. For 

Re=50,000, F=0.3 is sufficient to eliminate separation. It does not 

appear to matter if a particular frequency is achieved by reducing the 

spacing between wakes or by increasing the wake generator velocity, 

so long as F is sufficiently high. The required wake frequencies are 

about equal to the vortex generator jet pulsing frequencies required for 

separation control with steady inflow observed by Volino et al. [24, 

36]. 

s/Ls

t/T

0.55 0.6 0.65 0.7 0.75 0.8 0.85 0.9 0.95

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

0.5

1

1.5

2

2.5

3

s/Ls

t/T

0.6 0.65 0.7 0.75 0.8 0.85 0.9 0.95

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

1

2

3

4

5

6

Fig. 21 Phase averaged momentum thickness for low TI, 

Re=50,000: a) ζ=0.7, 2Lφ, b) ζ=1.4,1.6Lφ

CONCLUSIONS 

 The effect of unsteady wakes on the flow over the very high lift 

L1A airfoil was studied experimentally under low and high freestream 

turbulence conditions. Reynolds numbers based on suction surface 

length and nominal exit velocity of 25,000 and 50,000 were 

considered. Without wakes, the boundary layer separated and did not 

reattach. Wakes shed by upstream rods caused the boundary layer to 

reattach. When the wake passing frequency was low, either due to 

large spacing between wake generator rods or low rod velocity, the 

boundary layer separated between wakes. With a dimensionless wake 

passing frequency of F=0.5 or higher at Re=25,000 or with F=0.3 or 

higher at Re=50,000, separation was largely suppressed. Background 

freestream turbulence also helps to reduce separation, but the effect is 

secondary to the wake effect. 

 The expected wake passing frequency in an engine is near the 

frequency at which separation begins to be a problem on the L1A. 

The present cases will serve as baseline results for future cases in 

which wakes and flow control with vortex generator jets are combined. 

ACKNOWLEDGMENTS

 This work was sponsored by the National Aeronautics and Space 

Administration under grant NNC07IA10I. The grant monitor were 

Drs. Anthony Strazisar and James Heidmann of the NASA Glenn 

Research Center. The support of the United States Naval Academy 

Technical Support Department Shop and Fluids Laboratory is greatly 

appreciated. 

NASA/CR—2012-217415 121
12 

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No wakes

1.6Lφ, ζ=0.7

1Lφ, ζ=0.7

Inviscid

a)

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4

B5 B6

No wakes

1.6Lφ, ζ=0.7

1Lφ, ζ=0.7

b)

Fig. 22 Pressure profiles for high TI, Re=25,000 cases: a) 

Cp, b) Total pressure loss 

 

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φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

a)

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

b)

Fig. 23 Time space plot of phase averaged velocity 0.63Cx

downstream of cascade for high TI, Re=25,000, ζ=0.7, 1.6Lφ

rod spacing: a) U/Ue, b) u′/Ue

[9] Bons, J.P., Pluim, J., Gompertz, K., Bloxham, M., and Clark, J.P., 

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GT2008-50864. 

[10] Pluim, J., Memory, C., Bons, J., and Chen, J.P., 2009, “Designing 

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Reynolds Number and Periodic Unsteady Wake Flow Condition 

on Boundary Layer Development, Separation, and Intermittency 

Behavior Along the Suction Surface of a Low Pressure Turbine 

Blade,” ASME Journal of Turbomachinery, 129, pp. 92-107.

[12] Öztürk, B., and Schobeiri, M.T., 2007, “Effects of Turbulence 

Intensity and Periodic Unsteady Wake Flow Condition on 

Boundary Layer Development, Separation, and Reattachment 

Along the Suction Surface of a Low-Pressure Turbine Blade,” 

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[14] Mahallati, A., and Sjolander, S.A., 2007, “Aerodyanmics of a 

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Boundary Layer Development on an Ultra-High-Lift Low 

Pressure Turbine Airfoil,” ASME Journal of Turbomachinery, 

132, 011016. 

NASA/CR—2012-217415 122
13 

 

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

a)

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

b)

Fig. 24 Time space plot of phase averaged velocity 0.63Cx

downstream of cascade for high TI, Re=25,000, ζ=0.7, 1Lφ

rod spacing: a) U/Ue, b) u′/Ue

[16] Funazaki, K., Yamada, K., Ono, T., Segawa, K., Hamazaki, H., 

Takahashi, A., and Tanimitsu, H., 2006, “Experimental and 

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Variable Solidity,” ASME Paper GT2006-90507. 

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Surface Trips and Unsteady Wakes on the Boundary Layer 

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[20] Volino, R.J., 2010, “Separated Flow Measurements on a Highly 

Loaded Low-Pressure Turbine Airfoil,” ASME Journal of 

Turbomachinery, 132, 011007.

[21] Ibrahim, M., Kartuzova, O., and Volino, R.J., 2008, 

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IMECE2008-68879. 

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“Experimental and Computational Investigations of Separation 

and Transition on a Highly Loaded Low Pressure Turbine Airfoil: 

Part 2 – High Freestream Turbulence Intensity,” ASME Paper 

IMECE2008-68776.

 

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No wakes

1.6Lφ, ζ=0.7

1Lφ, ζ=0.7

1.6Lφ, ζ=1.4

Inviscid

a)

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4

B5 B6

No wakes

1.6Lφ, ζ=0.7

1Lφ, ζ=0.7

1.6Lφ, ζ=1.4

b)

Fig. 25 Pressure profiles for high TI, Re=50,000 cases: a) 

Cp, b) Total pressure loss 

[23] Volino, R.J., 2002, “Separated Flow Transition Under Simulated 

Low-Pressure Turbine Airfoil Conditions: Part 1 – Mean Flow 

and Turbulence Statistics,” ASME Journal of Turbomachinery, 

124, pp. 645-655. 

[24] Volino, R.J., Kartuzova, O., and Ibrahim, M.B., 2009, 

“Experimental and Computational Investigations of Low-Pressure 

Turbine Separation Control Using Vortex Generator Jets,” ASME 

Paper GT2009-59983. 

[25] Reimann, D., Bloxham, M., Pluim, J., and Bons, J.P., 2007, 

“Comparison of Spanwise Wake and Discrete Jet Disturbances on 

a Separating Low-Pressure Turbine Blade,” AIAA Paper 2007-

0525. 

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“Influence of Wake Passing Frequency and Elevated Turbulence 

Intensity on Transition,” AIAA Journal of Thermophysics and 

Heat Transfer, 19, 137-147. 

[27] Wills, J.A.B., 1962, “The Correction of Hot-Wire Readings for 

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Two-Dimensional Wakes Within Curved Channels: Theoretical 

Framework and Experimental Investigation,” ASME Journal of 

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[29] Pfeil, H. and Herbst, R., 1979, “Transition Procedure of 

Instationary Boundary Layers”, ASME Paper 79-GT-128. 

NASA/CR—2012-217415 123
14 

 

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

a)

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

b)

Fig. 26 Time space plot of phase averaged velocity 0.63Cx

downstream of cascade for high TI, Re=50,000, ζ=0.7, 1.6Lφ

rod spacing: a) U/Ue, b) u′/Ue

[30] Stieger, R. D., 2002, ‘‘The Effects of Wakes on Separating 

Boundary Layers in Low Pressure Turbines,’’ Ph.D. thesis, 

University of Cambridge, Cambridge, UK. 

[31] Stieger, R.D., and Hodson, H.P., 2005, “The Unsteady 

Development of a Turbulent Wake Through a Downstream LowPressure Turbine Blade Passage,” ASME Journal of 

Turbomachinery, 127, pp. 388-394. 

[32] Volino, R.J., 2005, “An Investigation of the Scales in Transitional 

Boundary Layers under High Freestream Turbulence 

Conditions,” Experiments in Fluids, 38, pp. 516-633. 

[33] Stieger, R.D., and Hodson, H.P., 2004, “The Transition 

Mechanism of Highly Loaded Low-Pressure Turbine Blades,” 

ASME Journal of Turbomachinery, 126, pp. 536-543. 

[34] Roberts, S.K., and Yaras, M.I., 2006, “Effects of SurfaceRoughness Geometry on Separation-Bubble Transition,” ASME 

Journal of Turbomachinery, 128, pp. 349-356. 

[35] Volino, R.J., and Bohl, D.G., 2004, “Separated Flow Transition 

Mechanisms and Prediction with High and Low Freestream 

Turbulence Under Low Pressure Turbine Conditions,” ASME 

Paper GT2004-53360. 

[36] Volino, R.J., Kartuzova, O., and Ibrahim, M.B., 2010, 

“Separation Control on a Very High Lift Low Pressure Turbine 

Airfoil Using Pulsed Vortex Generator Jets,” ASME Paper 

GT2010-23567. 

 

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

a)

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

b)

Fig. 27 Time space plot of phase averaged velocity 0.63Cx

downstream of cascade for high TI, Re=50,000, ζ=0.7, 1Lφ

rod spacing: a) U/Ue, b) u′/Ue

NASA/CR—2012-217415 124
1 Copyright © 2010 by ASME

Proceedings of ASME Turbo Expo 2010: Power for Land, Sea and Air

GT2009

June 14-18, 2010, Glasgow, UK

GT2010-23015

LES FLOW CONTROL SIMULATIONS FOR HIGHLY LOADED LOW PRESSURE 

TURBINE AIRFOIL (L1A) USING PULSED VORTEX GENERATOR JETS

Mounir B. Ibrahim, Olga Kartuzova

and Daniel J. Doucet

Mechanical Engineering Department

Cleveland State University

Cleveland, Ohio 44115-2425

m.ibrahim@csuohio.edu

Ralph J. Volino

Mechanical Engineering Department

United States Naval Academy

Annapolis, Maryland 21402-5042

volino@usna.edu

ABSTRACT 

Seven different cases were examined experimentally 

and computationally to study LPT flow control using pulsed 

VGJs for the L1A airfoil. These cases represent a combination 

of variation in Reynolds number, Re (25,000, 50,000 and 

100,000), based on the suction surface length and the nominal 

exit velocity from the cascade, blowing ratio, B (from 0.25 to 

1), dimensionless frequency, F (from 0.035 to 0.56) and duty 

cycle, DC (10% and 50%). The data was obtained for the 

pressure distribution along the airfoil and downstream in the 

wake as well as for velocity profiles at six different stations 

downstream of the suction peak. The CFD was done with LES 

utilizing version 6.3.26 of the finite-volume code ANSYS 

Fluent. The CFD did provide further insight to better 

understand the physics of flow control. The comparison 

between CFD and experimental results for Cp, velocity profiles 

and ψint is reasonable for all cases examined. 

Two of the cases examined did indicate that the higher 

DC could compensate for the lower F value. However, the 

effect of increasing the frequency appears to be stronger than 

increasing the DC value. The results from CFD using the QCriterion clearly illustrate how a separation bubble will persist 

at the lower frequency case (Case (2)) and the disturbances 

created from the jet flow do not have enough energy or time to 

travel further downstream to cause reattachment. On the other 

hand, the higher frequency case (Case (6)) did exhibit a 

penetration of the disturbance created by the jet into the 

separated region and caused reattachment, especially at the 

trailing edge. It appears that the jet was capable of breaking the 

large bubble into smaller ones with reattachment in between. 

NOMENCLATURE

B blowing ratio, maximum jet velocity/local freestream 

velocity

Cp ( ) 2

T − ρU/PP2 e , pressure coefficient

Cx axial chord length

DC duty cycle, fraction of time valve is open

F fLj-te/Uave, dimensionless frequency

f frequency

Lj-te distance from VGJs to trailing edge

Ls suction surface length

Lφ blade spacing (pitch)

P pressure

PS upstream static pressure

PT upstream stagnation pressure

PTe downstream stagnation pressure

NASA/CR—2012-217415 125
2 Copyright © 2010 by ASME

Re UeLs/ν, exit Reynolds number

s streamwise coordinate, distance from leading edge

T period of jet pulsing cycle

t time

u’ RMS of the fluctuating component of the streamwise 

velocity

TKE turbulent kinetic energy

Uave average freestream velocity between VGJs and trailing 

edge

Ue nominal exit freestream velocity, based on inviscid 

solution

Vx axial velocity

x axial distance from leading edge

ν kinematic viscosity

ρ density

ψ (PT-PTe)/(PT-PS), total pressure loss coefficient

ψint total pressure loss integrated over blade spacing

INTRODUCTION 

Boundary layer separation on the suction side of lowpressure turbine (LPT) airfoils can occur due to strong adverse 

pressure gradients. The problem is becoming more severe as 

airfoil loading is increased. If the boundary layer separates, the 

lift from the airfoil decreases and the aerodynamic loss 

increases, resulting in a drop in overall engine efficiency. A 

significant increase in efficiency could be achieved if separation 

could be prevented or minimized. Active flow control could 

provide a means for minimizing separation under conditions 

where it is most severe (low Re), without causing additional 

losses under other conditions (high Re). Minimizing separation 

will allow for improved designs with fewer stages and fewer 

airfoils per stage to generate the same power. The active flow 

control technique called Vortex Generator Jets (VGJs), 

introduced by Johnston and Nishi [1], is the subject of this 

paper. In this technique, blowing from small, compound angled 

holes is used to create streamwise vortices. The vortices bring 

high momentum fluid into the near wall region as well as 

promote transition and turbulent mixing. This helps to control 

separation. A review of experimental work on this topic was 

given in Volino et al. [2]. 

Along with the experimental investigations, numerical 

simulations of the flow over LPT blades, utilizing steady and 

pulsed vortex generator jets (VGJs), were performed by 

different researchers. This type of flow is challenging for CFD 

because of its transitional nature in combination with highly 

three-dimensional jet flow. 

Rizzetta and Visbal [3] used ILES (Implicit Large 

Eddy Simulation) to investigate the effect of flow control by 

pulsed VGJs on the flow separation in a Pack B cascade. The 

Pack B is a widely studied LPT airfoil. They reported that for 

inlet Re=25,000 and B=2, flow control helped to keep flow 

attached for an additional 15% of the chord. Despite some 

differences with experimental data, numerical and experimental 

time-mean velocity profiles were in reasonable agreement. 

Postl et al. [4] studied the effect of active flow control 

utilizing steady and pulsed VGJs on preventing laminar 

separation on the Pack B airfoil. In their study they used two 

computational approaches. The first was direct numerical 

simulation (DNS) of the flow over a flat plate with imposed 

streamwise pressure gradient as measured on the suction side of

the Pack B airfoil. In the second approach, a linear LPT cascade 

was simulated with the flow control cases performed in 2D 

using slots. They observed two different physical mechanisms 

to affect the control for the steady versus pulsed VGJs. The 

steady jets developed streamwise structures which enhanced 

momentum exchange between the near wall region and free 

stream. For the pulsed VGJs, an accelerated boundary layer 

transition helped to achieve flow reattachment. Their results are 

in reasonably good agreement with experimental data except in 

the separated region, where the size of the separation bubble 

was under predicted. 

Garg [5] used the NASA Glenn-HT code with the k-ω

SST model of Menter [6] to compute the flow over the Pack B 

blade with and without the use of VGJs. This work resulted in 

correct predictions of the separation location in the baseline 

case (without VGJs) as well as showing that separation vanishes 

in the flow control case as in the experiment. However, the 

separated region and the wake were not well predicted, which is 

common for RANS. 

In the present study the very highly loaded L1A airfoil 

was considered [7]. The L1A was designed at the Air Force 

Research Laboratory (AFRL) and is available on a limited basis 

from Clark [8]. It is an aft-loaded blade with the same flow 

angles and loading as the mid-chord loaded L1M. Based on the 

design calculations of Clark [8], the L1A has 10% higher 

loading than the “ultra-high lift” airfoils described by Zhang 

and Hodson [9], and 17% higher loading than the Pack B. 

Because the L1A is aft-loaded, it is more prone to separation 

than the L1M, as documented in Bons et al. [10], Ibrahim et al. 

[11], and Volino et al. [12]. In cases without flow control and 

with low freestream turbulence, the boundary layer separates 

when Re<150,000 and does not reattach, in spite of transition to 

turbulence in the shear layer over the separation bubble in all 

cases. This result contrasts with the results of earlier studies on 

less aggressive airfoils, which all showed reattachment after 

transition. The separation bubble on the L1A is about four 

times thicker than that on the Pack B. The larger distance from 

the shear layer to the wall on the L1A apparently prevents the 

turbulent mixing in the shear layer from reaching the wall and 

causing reattachment. The failure of the boundary layer to 

reattach results in a 20% loss in lift and increases profile losses 

by a factor of 7 [2]. At higher Reynolds numbers, the separation 

bubble is small and the boundary layer is attached over most of 

the airfoil. In cases with high freestream turbulence, results are 

similar, but the shear layer is somewhat thicker and the 

separation bubble thinner due to increased mixing induced in 

the shear layer. This results in reattachment after transition at 

Re=50,000 and 100,000. At the lowest Re considered (25,000) 

the boundary layer still does not reattach. 

NASA/CR—2012-217415 126
3 Copyright © 2010 by ASME

Boundary layer separation control has been studied 

using VGJs on the L1A airfoil. Experiments were done under 

low freestream turbulence conditions on a linear cascade in a 

low-speed wind tunnel (see Volino et al., [2]). In that paper, the 

pressure surveys on the airfoil surface and downstream total 

pressure loss surveys were documented. Cases were considered 

at Reynolds numbers (based on the suction surface length and 

the nominal exit velocity from the cascade) of 25,000, 50,000 

and 100,000. In all cases without flow control, the boundary 

layer separated and did not reattach. These cases (without jets) 

were documented experimentally and computationally for low 

freestream turbulence intensity (see Ibrahim et al. [12]) and 

high freestream turbulence intensity (see Volino et al. [13]). 

Cases with steady jets were documented for low freestream 

turbulence intensity (see Volino et al. [2] and Ibrahim et al. [7]).

Bons et al. [14] considered a case with Re=50,000, background 

freestream turbulence intensity of 3%, and periodic wakes 

produced with moving rods upstream of the airfoils. The VGJs 

had a duty cycle (fraction of time the jets are on during a 

pulsing cycle) of 30% and a blowing ratio of 2.3 (based on the 

maximum jet velocity in the pulsing cycle and the local 

freestream velocity). The dimensionless frequency of the wake 

passing and VGJ pulsing were both F= fLj-te/Uave,=0.34, where 

Lj-te is the streamwise distance from the VGJ holes to the trailing 

edge, and Uave, is the average freestream velocity over this 

distance. Two different streamwise locations for the VGJs were 

considered. The VGJs reduced the separation bubble size and 

reduced total pressure losses. 

This paper describes in detail the effects of using 

pulsed VGJs to control the flow over highly loaded LPT airfoil 

L1A discussed earlier. Experimental and computational 

methods are used to show how changes in pulsing frequency, 

blowing ratio, and duty cycle can achieve boundary layer 

separation control and even reattachment. 

 

EXPERIMENTAL FACILITY AND MEASUREMENTS 

Experiments were conducted in a closed loop wind 

tunnel with a seven blade linear cascade located in the wind 

tunnel’s third turn, as shown in Fig. 1a. Cascade parameters are 

shown in Table 1. A fine screen located upstream of the cascade 

is used to break up the boundary layers which form on the 

diffuser walls upstream of the test section and to provide 

uniform inlet conditions to the cascade. The freestream 

turbulence entering the cascade was measured with a cross-wire 

probe positioned just upstream of the center blade. The 

turbulence intensity is 0.8% in the streamwise component and 

0.5% in the cross stream components. The integral length scale 

of the streamwise component is 6.3 cm. Further details of the 

facility and inlet flow are in Volino [13]. 

A tailboard, shown in Fig. 1a, is needed to produce the 

correct exit flow angle from the cascade. Its position was set to 

produce periodicity at high Reynolds numbers. A tailboard on 

the opposite side of the cascade and inlet guide vanes were 

found to be unnecessary. To produce the correct approach flow 

to the end blades (B1 and B7), the amount of flow escaping 

around the two ends of the cascade is controlled with the flaps 

shown in Fig. 1a. The inlet flow angle at the center of the 

cascade was checked with a three-hole pressure probe and 

found to agree with the design angle to within 2° of uncertainty. 

Good periodicity at high Reynolds numbers was obtained in the 

exit flow. At low Reynolds numbers, when significant 

separation bubbles are present, the periodicity is not as good 

due to suppression of the separation bubble thickness on the 

blades closest to the tailboard.

Figure 1a. Schematic of linear cascade

Table 1: Cascade parameters

Axial 

Chord, Cx

[mm]

True 

Chord 

[mm]

Pitch,

Lφ

[mm]

Span

[mm]

Suction 

side, Ls

[mm]

Inlet 

flow 

angle

Exit 

flow 

angle

134 146 136 724 203 35° 60°

Figure 1b. Airfoil with VGJ holes and cross section of hole 

geometry.

In cases with effective flow control, periodicity is 

reestablished. The lack of periodicity in cases with large 

separation bubbles is considered acceptable since the focus of 

the study is separation control, and not documentation of cases 

NASA/CR—2012-217415 127
4 Copyright © 2010 by ASME

with large separation that would be unacceptable in practice. 

This compromise facilitates the study of a larger number of 

cases with flow control by obviating the need to adjust the 

tailboard by trial and error for each case. It also provides for 

better repeatability in the experiments, since the position of the 

tailboard is fixed for all cases. Any improvements made with 

flow control will be larger in practice than documented in the 

experiment, due to the effect of the tailboard in suppressing the 

bubble size in the uncontrolled cases. 

Each blade in the cascade has a central cavity which 

extends along the entire span. The cavity is closed at one end 

and has a fitting at the opposite end connected to a compressed 

air line. Air is supplied to the cavities from a common 

manifold. Manual ball valves are placed in the tubing between 

the manifold and blades to insure that each blade receives the 

same air flow. The valves also help to damp high frequency 

oscillations in the jet velocity when the VGJs are pulsed. The 

manifold is supplied through two fast response solenoid valves 

(Parker Hannifin 009-0339-900 with General Valve Iota One 

pulse driver) operating in parallel. The valves are supplied 

through a pressure regulator by the building air supply. A 

single spanwise row of holes was drilled into the suction surface 

of each blade at the inviscid pressure minimum location, 

s/Ls=0.5 (x/Cx=0.62), where s is the distance from the leading 

edge and Ls is the suction surface length. The pressure 

minimum has been shown in the studies listed above to be about 

the optimal location for flow control devices. The effects of 

devices located farther upstream are damped by the favorable 

pressure gradient, and devices located downstream of the 

separation point can also lose effectiveness. The holes are 0.8 

mm in diameter and drilled at 30° to the surface and 90° to the 

main flow direction (see Figure 1b). This is the same 

orientation used in all the VGJ studies listed above. The hole 

spacing is 10.6 diameters, and the length to diameter ratio is 12. 

The solenoid valves pulse the VGJs, and the pulsing 

frequency is presented below in dimensionless form as F= fLj-te/

Uave,, where Lj-te is the streamwise distance from the VGJ holes 

to the trailing edge, and Uave, is the average freestream velocity 

over this distance. For flow over single airfoils, F≥1 is typically 

needed to maintain separation control, but for cascades, Bons et 

al. [14] showed that control is possible in some cases with 

F=0.1. As shown in Volino [15] and Bons et al. [14], this is 

due to the extended calmed region which follows the jet 

disturbance. In practice, VGJs could be timed to wake passing 

in an LPT, which has a typical frequency of about F=0.3. 

The center blade, designated B4 in Fig. 1a, contains 

pressure taps near the spanwise centerline. Pressure surveys are 

made using a pressure transducer (0-870 Pa range Validyne 

transducer). Stagnation pressure is measured with a pitot tube 

upstream of the cascade. The uncertainty in the suction side 

pressure coefficients is 0.07. Most of this uncertainty is due to 

bias error. Stochastic error is minimized by averaging pressure 

transducer readings over a 10 second period. 

Total pressure losses are documented using a Kiel 

probe traversed across three blade spacings, 0.63 Cx

downstream of the cascade. A traverse is located in the wind 

tunnel downstream of the cascade to move the probe. The 

traverse causes an acceptably low blockage when it is located at 

least two Cx downstream of the cascade. 

Pressure and loss surveys were acquired at nominal 

Re=25,000, 50,000 and 100,000. The Reynolds number, as 

defined above, is based on the suction surface length and the 

nominal cascade exit velocity. The corresponding Reynolds 

numbers based on the cascade inlet velocity and the axial chord 

length are 10,000, 20,000 and 40,000, respectively. 

NUMERICAL SIMULATION 

The numerical simulations were conducted for the L1A

airfoil utilizing version 6.3.26 of the finite-volume code 

ANSYS Fluent [16]. In this paper, cases were considered at

different Reynolds numbers, blowing ratios, frequencies and 

duty cycles of blowing. Table 2 shows a summary of all the

CFD cases presented in this paper. The three dimensional 

computational domain includes a single passage (see Table 1 

for cascade parameters). A uniform velocity inflow condition is 

specified 1.9 Cx upstream of the blade leading edge in the flow 

direction. The inlet flow angle is set to 33o based on an inviscid 

calculation of the full cascade used in the experiment [11]. This 

angle agrees with the experimentally measured inlet angle to 

within the experimental uncertainty. The exit boundary is 

located 3.8 Cx downstream of the trailing edges in the flow 

direction. In the spanwise direction, the domain includes one 

VGJ. The boundary conditions on the sides of the passage are 

periodic. The full length of the feeding tube is included in the 

simulations, allowing the jet velocity profile to develop before 

entering the main domain [2]. A uniform velocity boundary 

condition during jet blowing is specified at the hole inlet, 

otherwise there would be no jet flow. For a duty cycle of 10%, 

this means the jet is on only for 10% of the cycle. Computation 

was continued until no variation (cycle-to-cycle) was reached. 

Convergence was established when residuals were 

reduced to a value 10-5 and the mass imbalance was less than 

0.01 %. 

Table 2. CFD Test matrix (NA = not available).

f, 

Hz

DC 

% 

Re/B

25,000/

1.0

50,000/

0.5

100,000/

0.25

3 10 Case (1) Case (2) Case (3) 

12 10 Case (4) Case (5) NA

24 10 NA Case (6) NA

12 50 NA Case (7) NA

NASA/CR—2012-217415 128
5 Copyright © 2010 by ASME

TURBULENCE MODEL DESCRIPTION 

Large Eddy Simulation with Dynamic Kinetic Energy 

Subgrid-Scale model 

In LES, large eddies are resolved directly, while small 

eddies are modeled. The rationale behind LES is as follows: a) 

Momentum, mass, energy, and other passive scalars are 

transported mostly by large eddies, b) Large eddies are more 

problem-dependent and are dictated by the geometries and 

boundary conditions of the flow involved, c) Small eddies are 

less dependent on the geometry, tend to be more isotropic, and 

are consequently more universal, and d) The chance of finding a 

universal turbulence model is much higher for small eddies.

The governing equations employed for LES are 

obtained by filtering the time-dependent Navier-Stokes 

equations in either Fourier (wave-number) space or 

configuration (physical) space. The filtering process effectively 

filters out the eddies whose scales are smaller than the filter 

width or grid spacing used in the computations. The resulting 

equations thus govern the dynamics of large eddies. The 

subgrid-scale stresses resulting from the filtering operation are 

unknown and require modeling. The subgrid-scale turbulence 

models in ANSYS Fluent employ the Boussinesq hypothesis as 

in the RANS models [16]. 

The dynamic subgrid-scale kinetic energy model used 

in present study is based on the model proposed by Kim and 

Menon [17]. In this model a separate transport equation is 

solved for subgrid-scale kinetic energy. The model constants 

are determined dynamically. The details of the implementation 

of this model in ANSYS Fluent and its validation are given by 

Kim [18]. 

RESULTS AND DISCUSSION 

Code Validation 

Three different grids were attempted for this study as 

shown in Table 3. Grid#3 was selected (as will be shown later) 

and Figure 2 shows this grid structure at: a) leading edge, b) 

trailing edge and c) jet exit, while Table 4 shows more 

specifications of the grid. 

 Table 3. Grids#1, 2 and 3 used in this investigation.

Grid 

# 

Size (Cells) Number 

of grids in 

z direction

y+ ∆z+ ∆x+

1 1,500,000 15 0.5 12.6 1 – 100

2 5,900,000 30 0.5 6.3 0.4 – 52

3 11,900,000 54 0.5 0.4 - 3.5 0.4 – 52

Table 4. Specification of computational Grid#3 

Number of cells 11.9 millions

Number of nodes on the suction surface 768

Number of nodes on the pressure surface 392

Number of nodes in span direction 54

y+ <1

∆z+ 0.4 - 3.5

∆x+ 0.4 - 52

Distance from inlet boundary to the 

leading edge

3.8 Cx

Distance from the trailing edge to the 

outlet boundary 

1.9 Cx

To accurately represent structures in the near-wall 

region (for LES), recommended values are: y+~ 2; ∆x+ ~ 50 -

150; ∆z+ ~ 15-40 (see Piomelli and Chasnov, [19]). Based on 

results (not shown for space limitation) for the pressure 

coefficient versus the dimensionless location on the suction side 

s/Ls , Grid#3 showed the closest agreement with the data and 

therefore was chosen for further computation. This grid was 

then run for different time steps (0.0005, 0.0001 and 0.00005 s) 

and a time step of 0.0001 s was selected since not much 

improvement was achieved using the smaller one (0.00005 s). 

Velocity measurements were conducted at six 

different stations downstream of the suction peak (see Table 5). 

Comparison will be made between the velocity profiles at these 

stations from CFD and experiments. 

Table 5. Velocity profile measurement stations.

Station 1 2 3 4 5 6

s/Ls 0.53 0.59 0.69 0.78 0.88 0.97

x/Cx 0.65 0.72 0.80 0.86 0.92 0.97

Pressure Distribution over Airfoil

Statistical averages of the pressure coefficient plotted 

versus dimensionless distance along the suction side of an 

airfoil are presented on Figures 3 through 9 for all (7) cases 

shown in Table 2. These cases represent a combination of 

variation in Re (25,000, 50,000 and 100,000), F (0.035 to 0.56) 

and duty cycle (10% and 50%). 

Figures 3, 4 and 5 show Cp for Cases (1) (F=0.14), (2) 

(F=0.07), and (3) (F=0.035), respectively with the same 

frequency (3 Hz) and duty cycle (10%). The Reynolds number 

varied from 25,000 to 100,000. The figures show flow 

separation starting after the suction peak with no reattachment, 

which is indicated by the large plateau in Cp after the suction 

peak. The experimental data are shown with error bar (0.07). Cp

at the suction peak is lower in CFD compared to experiment. 

The LES predictions are shown together with experimental data 

for the same case as well as no jet flow (due to the absence of 

experimental data for Cases (2) and (3)). It is clear that all cases 

exhibit separation with no reattachment similar to no jet 

blowing. The main reason for that is the low frequency at which 

NASA/CR—2012-217415 129
6 Copyright © 2010 by ASME

blowing occurred together with a low duty cycle that resulted in 

minimizing the effect of the jet in removing the bubble. 

Figures 6 and 7 show Cp for Cases (4) (F=0.56) and 

(5) (F=0.28) respectively, with the same frequency (12 Hz) and 

duty cycle (10%). The Reynolds number varied from 25,000 to 

50,000. The figures show flow separation starting after the 

suction peak with reattachment. The LES predictions are shown 

together with experimental data for the same case. The 

agreement between the CFD and experimental results is 

reasonable for most of the flow field. The main reason for the 

reattachment to occur is the higher frequency at which blowing 

occurred (compare Cases (1) and (4) and Cases (2) and (5)). 

Figure 8 shows Cp for Case (6) with Re=50,000, 

F=0.56 (f=24 Hz) and DC=10%. The figure shows flow 

separation starting after the suction peak with reattachment 

downstream as a result of the higher blowing frequency. The 

LES predictions are in reasonable agreement with experimental 

data for the same case. More details are given for Case (6) 

below. 

Figure 9 shows Cp for Case (7) with Re=50,000, 

F=0.28 (f=12 Hz) and DC=50%. The figure shows flow 

separation starting after the suction peak with reattachment. 

This case yielded similar results to Case (6) above. It is 

believed that increasing the duty cycle (from 10% to 50%) 

could result in the same effect as increasing the frequency from 

12 to 24 Hz. More discussion will be presented later on those 

effects. 

Predictions for Case (6) 

From the cases studied above, Case (6) was selected 

for more detailed examination. This case has experimental data 

for Cp and velocity profiles for comparison. 

Figure 10 shows the U velocity contours and velocity

vectors for Case (6). The flow shows a small separation bubble 

with reattachment. Velocity profiles normalized by the nominal 

exit velocity are plotted versus dimensionless distance from the 

wall in the direction normal to the wall on Fig. 11. The results 

for the six measurement stations (see Table 5) located 

downstream of the suction peak of the airfoil are shown. The 

agreement between the LES and experimental results is 

reasonable. Both CFD and experiment show separation starting 

at station 2 and a large separation bubble present at all stations 

from 2 to 5, based on negative velocities near the wall at those 

locations. However, the bubble size is larger in the experiment 

than in the CFD. It should be noted that the larger disagreement 

near the wall is due to the fact that hot wire anemometry was 

used in the experiment, which is not capable of measuring 

negative velocities. 

Profiles of the streamwise component of the RMS 

fluctuating velocity, u' (for both experiment and the resolved u’ 

from LES), normalized by the nominal exit velocity are plotted 

versus dimensionless distance from the wall in Fig. 12. The 

results for the six measurement stations located downstream of 

the suction peak of the airfoil are shown. LES results compare 

reasonably well with experiment in the magnitude of u'. 

Separation starts at station 2 and transition to turbulence starts 

between stations 3 and 4 (maximum subgrid TKE, u'). The 

location of the peak of u' is away from the wall (between station 

2 and 5) indicating the presence of a small bubble. The peak of 

u’ from the experiment is further away from the wall, indicating 

a larger bubble than predicted by CFD. 

For the purpose of visualization of the separated 

region as well as the influence of the jet's blowing ratio, isosurfaces of the mean axial velocity Vx =0.01 m/s are shown in 

Fig. 13 for Case (6). The reason for choosing Vx =0.01 m/s is 

that this small (but not negative) value represents velocity in the 

shear layer of the separation bubble and helps to visualize the 

size of the bubble and shapes of the vortices created by the jets. 

The airfoil with three jets on the suction side near the suction 

peak is shown. The visualization shows that mixing is 

happening near the wall which causes the flow to reattach near 

the trailing edge. 

 

Velocity Contours at Jet Exit 

CFD analysis provides further insight into the physics 

of the problem compared to the experiments. The velocity 

contours were examined at the jet exit for Cases (2), (5), and 

(6). Those cases have the same Re (50,000), B (0.5) and DC 

(10%) values but vary in F (0.07, 0.28 and 0.56 for the three 

cases respectively). Figures 14, 15 and 16 show the contours of 

the velocity magnitude at the jet exit for the three cases 

respectively. The traveling time in the feeding tube for all cases 

is about 6.1 ms. The blowing times, however, vary for the three 

cases (33.3 ms for Case (2); 8.33 ms for Case (5); and 4.17 ms 

for Case (6)). A lower frequency results in a higher blowing 

time. Thus more time is given for the flow to reach the jet exit 

and exhibit a parabolic profile (compare Figures 14 and 16). On 

the other hand, lower frequencies (with the same DC) result in 

the higher quiet time (no blowing) and thus the case gets closer 

to uncontrolled one. This explains the poor flow control results 

(flow separation) in Case (2) as shown earlier in Figure 4. 

Figure 17 shows the velocity contours at the jet exit 

for Case (7). This case has Re=50,000, B=0.5, DC=50% and 

F=0.28. One additional feature that exists in this case (that 

differs from the above Cases (2), (5) and (6)) is the fact that the 

blowing during the 50% DC was split into two parts. The first 

10% was at the nominal blowing value while that velocity was 

reduced to lower values at the second part of the blowing. This 

was done to match the jet velocity profile seen in the 

experiment with no cross flow present (see Figure 2 of Volino 

et al. [2]). 

 

Effect of Re, F and DC 

Table 6 shows the main characteristics for all the cases 

examined. The seven cases represent a combination of the 

variation in Re, F and DC. The total pressure loss integrated 

over blade spacing, ψint , is also shown in the table for both CFD 

and experimental results. 

NASA/CR—2012-217415 130
7 Copyright © 2010 by ASME

Table 6. Main Parameters of all Cases Examined. 

(NA = not available)

Case 1 2 3 4 5 6 7 

Re/1000 25 50 100 25 50 50 50

B 1.0 0.5 0.25 1.0 0.5 0.5 0.5

DC % 10 10 10 10 10 10 50

f, Hz 3 3 3 12 12 24 12

Uave,, 

m/s

2.17 4.35 8.7 2.17 4.35 4.35 4.35

F 0.14 0.07 0.035 0.56 0.28 0.56 0.28

ψint,

CFD

0.923 1.026 0.825 0.515 0.372 0.246 0.384

ψint Exp NA NA NA 0.346 0.356 0.237 0.313

As the Reynolds number increases the losses decrease. 

Cases (4) and (6) show that despite having the same F value, 

more losses are encountered at lower Re. 

As the value of F increases the losses decrease, see 

Cases (2), (5) and (6). It appears that a value of F of 0.28 or 

above is sufficient enough to cause reattachment and possible 

removal of the bubble, depending on the Reynolds number. 

Cases (6) and (7) show that the higher DC (Case (6)) 

could compensate for the lower F (Case (7)). However, the 

effect of increasing the frequency appears to be stronger than 

increasing the DC value. 

The comparison between CFD and experimental 

results for ψint is good for Cases (5), (6) and (7). As for Case (4) 

the larger differences is due to the fact that it is not a fully 

attached case and therefore there is no complete periodicity 

across the cascade in the experiment. Therefore, the 

experimental value is expected to be lower than the CFD 

equivalent. 

 

Q-Criterion Applied for Cases (2) and (6) 

In order to show the effect of frequency, the QCriterion was used for two cases: Case (2)-low frequency, and 

Case (6)-high frequency. The Q-Criterion is defined as 

Q=1⁄2(ΩijΩij-SijSij), which is the second invariant of the velocity 

gradient tensor. The components Ωij and Sij are the symmetric 

and antisymmetric parts of the velocity gradient tensor. 

For the case of Re=50,000 and B=0.5 under steady 

blowing a separation bubble will exist on the airfoil [7]. 

Furthermore, at the lower frequency (Case (2)), the timeaveraged flow field shows separation without reattachment (see 

Figure 4). At the higher frequency (Case (6)), the time-averaged 

flow field shows separation and reattachment (see Figure 8). 

Figure 18 shows the iso-surfaces of Q-Criterion 

colored by Vx (m/s) for Case (2) at different times in the cycle. 

At the beginning of blowing, a large separation bubble is 

present and the shear layer above the jet is undisturbed. In the 

middle of blowing, the large separated region remains above the 

airfoil surface but the boundary layer downstream of the jet is 

energized. Shortly after shut down (t/T=10% of the cycle), the 

boundary layer in the vicinity of the jet starts to relax with the 

energized region moving downstream. This energized region 

causes shrinkage of the separation bubble downstream of the 

jet, but it does not have enough energy or time to travel further 

downstream to cause reattachment. At t/T=80% of the cycle 

after the jet’s shutdown, the flow looks very similar to steady 

blowing where a large separation bubble does exist as indicated 

above. 

Figure 19 shows the iso-surfaces of Q-Criterion 

colored by Vx (m/s) for Case (6) at different times in the cycle. 

At the beginning of blowing there is a large separation bubble 

present near the trailing edge of the airfoil. During blowing the 

separated region is traveling downstream and the flow becomes 

attached in that region. Right after shutdown of the jet 

(t/T=10% of the cycle time) an overall smaller separation 

region is present and the flow starts to reattach at the trailing 

edge. At t/T=80% of the cycle after the jet’s shutdown, the flow 

is attached at the trailing edge, but the separated region starts to 

appear upstream. 

 

SUMMARY AND CONCLUSIONS 

Seven different cases were examined experimentally 

and computationally to study LPT flow control using pulsed 

VGJs for the L1A airfoil. These cases represent a combination 

of variation in Re (25,000, 50,000 and 100,000), based on the 

suction surface length and the nominal exit velocity from the 

cascade, B (from 0.25 to 1.0), F (from 0.035 to 0.56) and DC

(10% and 50%). The data was obtained for the pressure 

distribution along the airfoil and downstream in the wake as 

well as velocity profiles at six different stations downstream of 

the suction peak. The CFD was done with LES utilizing version 

6.3.26 of the finite-volume code ANSYS Fluent. The CFD 

provided further insight to better understand the physics of the 

flow control. 

All cases examined did show flow separation with no 

jet blowing. At lower Re, a larger separation bubble existed and 

accordingly it was more difficult to remove. As the Re 

increased, the separation bubble size was reduced and the losses 

decreased. As the value of F increased, the losses decreased and 

it appears that a value of F>0.28 is sufficient for achieving 

reattachment and possible removal of the bubble, depending on 

the Re. Two of the cases examined did indicate that a higher 

DC could compensate for the lower F value. However, the 

effect of increasing the frequency appeared to be stronger than 

increasing the duty cycle value. The comparison between CFD 

and experimental results for Cp, velocity profiles and ψint are 

reasonably good for all cases examined. The LES predictions 

show smaller separation bubbles than those measured in the 

experiment. This could be attributed to : 1) the mesh coarsening 

in the normal to wall direction was not fine enough and/or 2) 

the inadequate turbulence model used in the subgrid to resolve

the turbulence near the wall.

The Q-Criterion was used for two cases: Case (2)-low 

frequency, and Case (6)-high frequency. The results of CFD 

clearly illustrate how a separation bubble will persist in the 

NASA/CR—2012-217415 131
8 Copyright © 2010 by ASME

lower frequency case and the disturbances created from the jet 

flow do not have enough energy or time to travel further 

downstream to cause reattachment. On the other hand, the 

higher frequency case did exhibit a penetration of the 

disturbance created by the jet into the separated region and 

cause reattachment, especially at the trailing edge. It appears 

that the jet was capable of breaking the large bubble into 

smaller ones, with reattachment in between. 

ACKNOWLEDGMENTS 

This work was sponsored by the National Aeronautics 

and Space Administration. The grant monitor is Dr. Anthony 

Strazisar of the NASA Glenn Research Center. The support of 

the United States Naval Academy Technical Support 

Department Shop and Fluids Laboratory is greatly appreciated. 

We also greatly appreciate the computer time provided for us 

by the Ohio Super Computer (OSC). The OSC Computer 

Cluster has been made available as part of the center’s mission 

to support Ohio universities.

REFERENCES

1. Johnston, J.P. and Nishi, M. “Vortex Generator Jets. Means for 

Flow Separation Control”, AIAA Journal, 28, pp. 989-994, 

1990.

2. Volino, R.J., Kartuzova, O. and Ibrahim, M.B. “Experimental 

and Computational Investigations of Low-Pressure Turbine 

Separation Control using Vortex Generator Jets”, GT2009-

59983, Proceedings of ASME Turbo Expo 2009: Power for 

Land, Sea and Air GT2009, June 8-12, 2009, Orlando, Florida, 

USA. 

3. Rizzetta, D.P. and Visbal, M.R., “Numerical Simulation of 

Separation Control for Transitional Highly Loaded LowPressure Turbines”, AIAA Journal, 43, pp.1958-1967, 2005.

4. Postl, D., Gross, A., Fasel, H.F. “Numerical Investigation of 

Active Control for Low-Pressure Turbine Blade Separation”,

AIAA2004-750, Proceedings of 42nd AIAA Aerospace Science 

Meeting and Exhibit, January 5-8, 2004, Reno, Nevada, USA, 

2004.

5. Garg, V.K., “Low-Pressure Turbine Separation Control - 

Comparison with Experimental Data”, ASME Paper GT-2002-

30229, 2002.

6. Menter, F.R. “Two-Equation Eddy-Viscosity Turbulence 

Models for Engineering Applications”, AIAA Journal, 32, pp. 

1598-1605, 1994.

7. Ibrahim, M.B., Kartuzova, O., and Volino, R.J., “LES and 

URANS Computational Investigations of LPT Blade (L1A) 

Separation Control using Vortex Generator Jets”, Proceedings of 

Turbulence, Heat and Mass Transfer 6, 2009, Rome, Italy.

8. Clark, J.P. Private Communication, Air Force Research 

Laboratory, 2007.

9. Zhang, X.F. and Hodson, H. “Combined Effects of Surface Trips 

and Unsteady Wakes on the Boundary Layer Development of an 

Ultra-High-Lift LP Turbine Blade”, ASME Journal of 

Turbomachinery, 127, pp. 479-488, 2005.

10. Bons, J.P., Pluim, J., Gompertz, K. and Bloxham, M. “The 

Application of Flow Control to an Aft-Loaded Low Pressure 

Turbine Cascade with Unsteady Wakes”, ASME Paper GT2008-

50864, 2008.

11. Ibrahim, M.B., Kartuzova, O. and Volino, R.J., “Experimental 

and Computational Investigations of Separation and Transition 

on a Highly Loaded Low Pressure Turbine Airfoil: Part 1 – Low 

Freestream Turbulence Intensity”, ASME Paper IMECE2008-

68879, 2008.

12. Volino, R.J., Kartuzova, O. and Ibrahim, M.B., “Experimental 

and Computational Investigations of Separation and Transition 

on a Highly Loaded Low Pressure Turbine Airfoil: Part 2 – High 

Freestream Turbulence Intensity”, ASME Paper IMECE2008-

68776, 2008.

13. Volino, R.J., 2008, “Separated Flow Measurements on a Highly 

Loaded Low-Pressure Turbine Airfoil,” ASME Paper GT2008-

51445.

14. Bons, J.P., Sondergaard, R., and Rivir, R.B., 2002, “The Fluid 

Dynamics of LPT Blade Separation Control Using Pulsed Jets,” 

ASME Journal of Turbomachinery, 124, pp. 77-85.

15. Volino, R.J., 2003, “Separation Control on Low-Pressure 

Turbine Airfoils Using Synthetic Vortex Generator Jets,” ASME 

Journal of Turbomachinery, 125, pp. 765-777.

16. ANSYS Fluent, Inc., 2005. Fluent 6.3 - User Guide.

17. Kim, W.W. and Menon, S. “Application of the localized 

dynamic subgrid-scale model to turbulent wall-bounded flows”, 

Technical Report AIAA-97-0210, American Institute of 

Aeronautics and Astronautics, 35th Aerospace Sciences 

Meeting, Reno, NV, 1997.

18. Kim, S.E. “Large eddy simulation using unstructured meshes 

and dynamic subgridscale turbulence models”, Technical Report 

AIAA-2004-2548, American Institute of Aeronautics and 

Astronautics, 34th Fluid Dynamics Conference and Exhibit, 

2004.

19. Piomelli, U. and Chasnov, J.R., “Large-Eddy Simulations: 

Theory and Applications”, Turbulence and Transition Modeling, 

lecture notes from the ERCOFTAC/IUTAM summer school held 

in Stockholm, 12-20 June, 1995.

NASA/CR—2012-217415 132
9 Copyright © 2010 by ASME

a) Leading Edge

b) Trailing Edge

c) Jet Exit

Re = 25,000

0.0

0.5

1.0

1.5

2.0

2.5

0 0.2 0.4 0.6 0.8 1

s/Ls

Cp

Exp, B=1, f=3, DC=10

Exp., No Jet

CFD, B=1, f=3, DC=10

Figure 3. Cp for Case (1), Re=25,000, B=1.0, F=0.14, 

DC=10%. 

Re = 50,000

0.0

0.2

0.4

0.6

0.8

1.0

1.2

1.4

1.6

1.8

2.0

0 0.2 0.4 0.6 0.8 1 s/Ls

Cp

Exp, No Jet

CFD, B=0.5, f=3, DC=10

Figure 4. Cp for Case (2), Re= 50,000-B=0.5-F=0.07, 

DC=10%.

Re = 100,000

0.0

0.5

1.0

1.5

2.0

2.5

0 0.2 0.4 0.6 0.8 1 s/Ls

Cp

Exp, No Jet

CFD, B=0.25, f=3, DC=10

Figure 5. Cp for Case (3), Re= 100,000, B=0.25, 

F=0.035, DC=10%.

Figure 2. Grid#3 used in pulsed jets flow control 

cases: a) Leading Edge, b) Trailing Edge and c) 

Jet Exit.

NASA/CR—2012-217415 133
10 Copyright © 2010 by ASME

Re = 25,000

0.0

0.5

1.0

1.5

2.0

2.5

0 0.2 0.4 0.6 0.8 1

s/Ls

Cp

Exp, B=1, f=12, DC=10

CFD, B=1, f=12, DC=10

Figure 6. Cp for Case (4), Re=25,000, B=1, 

F=0.56, DC=10%.

Re = 50,000

0.0

0.5

1.0

1.5

2.0

2.5

3.0

0 0.2 0.4 0.6 0.8 1

s/Ls

Cp

CFD, B=0.5, f=12, DC=50

Exp., B=0.5, f=12, DC=50

Exp., B=1, f=12, DC=10

Figure 9. Cp for Case (7), Re=2, 50,000, B=0.5, F=0.28, DC=50%.

Re = 50,000

0.0

0.5

1.0

1.5

2.0

2.5

0 0.2 0.4 0.6 0.8 1 s/Ls

Cp

Exp. B=1, f=12 Hz, DC=10

Exp, B=0.5, f=12, DC=10

CFD, B=0.5, f=12, DC=10

Figure 7. Cp for Case (5), Re= 50,000, B=0.5,

F=0.28, DC=10%. 

Figure 10. Contours of U, m/s, and velocity vectors 

for Case (6), Re=50,000, B=0.5, F=0.56, DC=10%.

Re = 50,000

0.0

0.5

1.0

1.5

2.0

2.5

0.00 0.20 0.40 0.60 0.80 1.00

s/Ls

Cp

Exp, B=0.5, f=24, DC=10

CFD, B=0.5, f=24, DC=10

Figure 8. Cp for Case (6), Re=50,000, B=0.5, F=0.56, 

DC=10%.

NASA/CR—2012-217415 134
11 Copyright © 2010 by ASME

s/Ls = 0.53

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0 0.5 1 1.5

U/Ue

y/Ls

Exp., B=0.5, f=24 Hz, DC=10

CFD, B=0.5, f=24 Hz, DC=10 s/Ls = 0.78

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

-0.5 0 0.5 1 1.5

U/Ue

y/Ls

s/Ls = 0.59

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

-0.5 0 0.5 1 1.5

U/Ue

y/Ls

s/Ls = 0.88

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

-0.5 0 0.5 1 1.5

U/Ue

y/Ls

s/Ls = 0.69

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

-1 0 1 2

U/Ue

y/Ls

s/Ls = 0.97

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0 0.5 1 1.5

U/Ue

y/Ls

Figure 11. U/Ue at the different 6 stations for Case (6), Re=50,000, 

B=0.5, F=0.56, DC=10%.

NASA/CR—2012-217415 135
12 Copyright © 2010 by ASME

s/Ls = 0.53

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0 0.05 0.1 0.15

u'/Ue

y/Ls

Exp., B=0.5, f=24 Hz, DC=10

CFD, B=0.5, f=24 Hz, DC=10 s/Ls = 0.78

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0 0.2 0.4 0.6

u'/Ue

y/Ls

s/Ls = 0.59

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0 0.05 0.1 0.15 0.2

u'/Ue

y/Ls

s/Ls = 0.88

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0 0.2 0.4 0.6

u'/Ue

y/Ls

s/Ls = 0.69

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0 0.1 0.2 0.3

u'/Ue

y/Ls

s/Ls = 0.97

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0 0.1 0.2 0.3 0.4

u'/Ue

y/Ls

Figure 12. u’/Ue at the different 6 stations for Case (6), 

Re=50,000, B=0.5, F=0.56, DC=10%.

NASA/CR—2012-217415 136
13 Copyright © 2010 by ASME

Figure 13. Isosurface of mean Vx = 0.01 m/s

for Case (6), Re=50,000, B=0.5, F=0.56, DC=10%. 

Figure 14. Velocity 

contour (m/s) at the jet 

exit for Case (2)

Re=50,000, B=0.5, 

F=0.07, DC=10%.

Figure 15. Velocity 

contour (m/s) at the jet 

exit for Case (5)

Re=50,000, B=0.5, 

F=0.28, DC=10%.

NASA/CR—2012-217415 137
14 Copyright © 2010 by ASME

Figure 16. Velocity 

contour (m/s) at the jet 

exit for Case (6)

Re=50,000, B=0.5, 

F=0.56, DC=10%.

Figure 17. Velocity 

contour (m/s) at the jet 

exit for Case (7)

Re=50,000, B=0.5, 

F=0.28, DC=50%.

NASA/CR—2012-217415 138
15 Copyright © 2010 by ASME

Figure 18. Q-criterion contours colored by Vx (m/s), 

Cases (2), Re= 50,000, B=0.5, F=0.07, DC=50%, at 

different times in the cycle.

Figure 19. Q-criterion contours colored by Vx (m/s), 

Cases (6), Re= 50,000, B=0.5, F=0.56, DC=10%, at 

different times in the cycle.

NASA/CR—2012-217415 139
1 

This material is declared a work of the U.S. Government and is

Proceedings of ASME Turbo Expo 2011 

GT2011 

June 6-10, 2011, Vancouver, Canada

GT2011-46106

COMBINED EFFECTS OF WAKES AND PULSED VORTEX GENERATOR JET FLOW 

CONTROL ON BOUNDARY LAYER SEPARATION ON A VERY HIGH LIFT LOW 

PRESSURE TURBINE AIRFOIL 

Ralph J. Volino

Mechanical Engineering Department 

United States Naval Academy 

Annapolis, Maryland 21402-5042 

volino@usna.edu 

ABSTRACT 

 Boundary layer separation control with pulsed vortex generator jets 

(VGJs) has been studied on a very high lift, low-pressure turbine 

airfoil in the presence of unsteady wakes. Experiments were done 

under low (0.6%) and high (4%) freestream turbulence conditions on a 

linear cascade in a low speed wind tunnel. Cases were considered at 

Reynolds numbers (based on the suction surface length and the 

nominal exit velocity from the cascade) of 25,000 and 50,000. Wakes 

were produced from moving rods upstream of the cascade with flow 

coefficient 1.13 and rod spacing equal 2 blade pitches, resulting in a 

dimensionless wake passing frequency F=fLj-te/Uave=0.14, where f is 

the frequency, Lj-te is the length of the adverse pressure gradient region 

on the suction surface, and Uave is the average freestream velocity. 

The VGJs were injected at the beginning of the adverse pressure 

gradient region on the suction surface with maximum jet velocity in 

each pulse equal to the local freestream velocity and a jet duty cycle of 

10%. Several different timings of the VGJs with respect to the wakes 

were considered. Pressure surveys on the airfoil surface and 

downstream total pressure loss surveys were documented. 

Instantaneous velocity profile measurements were acquired in the 

suction surface boundary layer and downstream of the cascade. In 

cases without VGJs, the boundary layer momentarily reattached in 

response to the wake passing, but separated between wakes. The 

VGJs also caused reattachment, and if the VGJ pulsing frequency was 

sufficiently high, separation was largely suppressed for the full wake 

passing cycle. The timing of the VGJs with respect to the wakes was 

not very important. The jet pulsing frequency needed for separation 

control was about the same as found previously in cases without 

wakes. The background freestream turbulence effect was negligible in 

the presence of the larger wake and VGJ disturbances. 

NOMENCLATURE 

Cp 2(PT-P)/(ρUe

2

), pressure coefficient 

Cx axial chord length 

F fLj-te/Uave, dimensionless frequency 

f wake passing frequency 

Lj-te length of adverse pressure gradient region on suction surface

Ls suction surface length 

Lφ blade spacing (pitch) 

P pressure 

PS upstream static pressure 

PT upstream stagnation pressure 

PTe downstream stagnation pressure 

Re UeLs/ν, exit Reynolds number 

s streamwise coordinate, distance from leading edge 

T period of jet pulsing cycle 

t time 

TI background freestream turbulence intensity 

U local mean velocity 

Uave average freestream velocity in adverse pressure gradient region 

Ui inlet freestream velocity 

Ue nominal exit freestream velocity, based on inviscid solution 

Urod wake generator velocity 

u′ rms fluctuating streamwise velocity 

x axial distance from leading edge 

αi inlet flow angle 

φ coordinate along blade spacing, normal to axial chord 

ν kinematic viscosity 

ρ density 

ψ (PT-PTe)/(PT-PS), total pressure loss coefficient 

ζ Uicos(αi)/Urod = Uaxial/Urod, flow coefficient

INTRODUCTION 

 Boundary layer separation on the suction side of low-pressure 

turbine (LPT) airfoils can cause partial loss of lift and high 

aerodynamic losses (Hourmouziadis [1], Mayle [2], Sharma et al. [3]). 

In aircraft engines the lower Reynolds numbers at altitude can lead to 

a component efficiency drop of 2% between takeoff and cruise in large 

commercial transport engines, and possibly as much as 7% in smaller 

engines operating at higher altitudes [4, 5]. Separation becomes more 

likely when airfoil loading is high because of the strong adverse 

pressure gradients on the suction surface, but high loading is desirable 

since it can be used to reduce airfoil count, weight and cost. Accurate 

prediction of separation under relevant conditions, including the 

effects of boundary layer transition and periodic unsteadiness, is 

needed to design high lift airfoils without separation problems. 

Proceedings of ASME Turbo Expo 2011 

GT2011

June 6-10, 2011, Vancouver, British Columbia, Canada

GT2011-4

1

NASA/CR—2012-217415 140 This material 

distribution is unlimited. 

is declared a work of the U.S. Government and is not subject to copyright protection in the United States. 

Approved for public release. Distribution is unlimited.
2 

This material is declared a work of the U.S. Government and is

 not subject to copyright protection in the United States. Approv

distribution is unlimited. ed for public release; 

 Separation can be mitigated in a few ways. One is by wakes shed 

from the airfoils in upstream stages in an engine. The velocity deficit 

and elevated turbulence in periodic wakes help to suppress separation 

and can cause a separated boundary layer to reattach. Hodson and 

Howell [6] describe the mechanisms by which wakes promote 

reattachment, including the “negative jet” which results when the 

velocity deficit in the wake causes the flow outside the wake to 

accelerate and impinge on the surface, and the unsteadiness which 

promotes transition in the boundary layer. Following the wake itself is 

a calmed period (Gostelow et al. [7] and Schulte and Hodson [8]) in 

which the boundary layer has low turbulence and greater resistance to 

separation. Numerous studies have considered the wake effect in the 

LPT, including those listed in Hodson and Howell [6], and more recent 

references in Bons et al. [9] and Pluim et al. [10]. Examples include 

Schobeiri et al. [11], Öztürk and Schobeiri [12], Jiang and Simon [13], 

and Mahallati and Sjolander [14] who all used the Pack B airfoil. 

Zhang and Hodson [15] and Funazaki et al. [16] used more highly 

loaded airfoils. Many additional studies are available from these 

research groups and others. 

 Separation problems can also be limited through good airfoil design, 

as described by Praisner and Clark [17]. In recent years, knowledge of 

wake effects has allowed for designs with higher loading than would 

be possible under steady inflow conditions. Even with the best design 

methods, however, a loading limit will always exist, above which 

separation will still occur. Flow control, either active or passive, 

might allow an extension of this limit. 

 Separation control with passive devices such as boundary layer trips 

has been shown effective by Zhang et al. [18], Bohl and Volino [19], 

Volino [20], and others. Passive devices have the distinct advantage 

of simplicity, but they also introduce parasitic losses and cannot be 

adjusted to account for changes in flow conditions. Active devices are 

also possible, and although their complexity and reliability would 

create challenges, they could be made adjustable and provide 

potentially better control. In turbomachinery, plasma devices as used 

by Huang et al. [21] could be viable, and are under active study. 

Vortex generator jets (VGJs), as introduced by Johnston and Nishi 

[22], have also been considered. Blowing from small, compound 

angled holes is used to create streamwise vortices which promote 

transition and bring high momentum fluid into the near wall region to 

help control separation. The most effective VGJs enter the boundary 

layer at a relatively shallow pitch angle (typically 30 to 45 degrees) 

relative to the wall and a high skew angle (45 to 90 degrees) relative to 

the main flow. Bons et al. [4, 23], Volino [24], Volino and Bohl [25], 

McQuilling and Jacob [26], and Eldredge and Bons [27] all used VGJs 

on the highly loaded Pack B LPT airfoil. Separation was essentially 

eliminated, even at the lowest Reynolds number considered. Similar 

results with were found on the very highly loaded L1M airfoil by Bons 

et al. [28], who saw the size of a large separation bubble reduced by 

VGJs. Pulsed jets were more effective than steady jets in all studies. 

The initial disturbance created by each pulse caused the boundary 

layer to attach. The boundary then remained resistant to separation 

during the calmed period which followed the VGJ disturbance. When 

the time between pulses was long enough, the boundary layer did 

eventually relax to a separated state, but due to the control which 

persisted during the calmed period, the VGJs were effective even with 

low jet pulsing frequencies, duty cycles and mass flow rates. Since the 

boundary layer was attached and undisturbed for much of the jet 

pulsing cycle, profile losses were low. 

 The present study uses the L1A airfoil, which was designed at the 

Air Force Research Laboratory (AFRL) and is available on a limited 

basis from Clark [29]. Dimensions of the L1A as used in the present 

study are given in Table 1. It was deliberately designed to provide a 

challenging case for flow control. The L1A has a Zweifel coefficient 

of 1.35, which corresponds to 10% higher loading than the “ultra-high 

lift” airfoils described by Zhang and Hodson [30], and 17% higher 

loading than the Pack B airfoil. The L1A is also aft loaded, which is 

advantageous for reducing secondary flow losses at the endwalls, but 

makes the boundary layer more prone to separation than a forward 

loaded blade, as documented in Bons et al. [9], Volino [31], Ibrahim et 

al. [32], and Volino et al. [33]. In cases without wakes and low 

Reynolds numbers, the boundary layer separates and does not reattach, 

in spite of transition to turbulence in the shear layer over the 

separation bubble. This result contrasts with the results of studies on 

less aggressive airfoils (e.g. Volino [34]), which all showed 

reattachment after transition. The failure to reattach can occur even 

when transition starts farther upstream on the L1A than on other 

airfoils (e.g. with low freestream turbulence and Re=100,000, 

transition starts at s/Ls=0.8 on the Pack B and causes reattachment, and 

at s/Ls=0.6 on the L1A without reattachment). The adverse pressure 

gradient on the L1A is roughly twice as strong as on the Pack B, and is 

apparently strong enough to prevent reattachment at low Reynolds 

numbers in spite of transition and turbulent mixing in the shear layer 

over the separation bubble. The failure of the boundary layer to 

reattach results in a 20% loss in lift and increases profile losses by up 

to a factor of 7. At higher Reynolds numbers the separation bubble 

closes, and for Re 200,000 the separation bubble on the L1A is small 

and the boundary layer is attached over most of the airfoil. 

 Two studies have considered the effect of wakes on the L1A 

boundary layer. Bons et al. [9] considered a case with Re=50,000 

(based on the suction surface length and the nominal exit velocity from 

the cascade), background freestream turbulence TI=3%, and periodic 

wakes produced with moving rods upstream of the airfoils. The 

dimensionless frequency of the wake passing was F=fLj-te/Uave=0.34, 

where Lj-te is the length of the adverse pressure gradient region on the 

suction surface, and Uave is the average freestream velocity over this 

distance. The length Lj-te is also the distance from a row of vortex 

generator jet (VGJ) holes to the trailing edge. Volino [35] considered 

cases at high (4%) and low (0.6%) freestream turbulence with 

Re=25,000 and 50,000. The spacing and speed of moving rods were 

varied to produce wake passing frequencies between F=0.14 and 0.56. 

Wakes largely suppressed separation at Re=25,000 when F was above 

0.5. At lower frequencies the disturbances caused by the wakes 

caused momentary reattachment, but the boundary layer re-separated 

between wake passing events. For Re=50,000, F=0.3 was sufficient to 

largely suppress separation. The effect was the same whether a 

particular frequency was achieved by changing rod spacing or rod 

velocity. Higher freestream turbulence helped to promote transition 

and reattachment, but the effect was small compared to the wake 

passing effect. 

 Flow control with vortex generator jets on the L1A has been 

considered in Bons et al. [9], Volino et al. [36, 37, 38], and Ibrahim et 

al. [39, 40]. The same Reynolds numbers and freestream turbulence 

levels were considered as with the wakes. With a VGJ blowing ratio 

of 1 (i.e. maximum jet velocity equal to the freestream velocity) and 

10% duty cycle, a dimensionless jet pulsing frequency of F=0.5 was 

sufficient to control separation with Re=25,000, and F=0.3 was 

sufficient for Re=50,000. These frequencies match the wake passing 

frequencies required for separation control at each Reynolds number, 

suggesting that the type of disturbance is not as important as the 

frequency of the disturbance for controlling the boundary layer. At 

lower frequencies the flow control was not as good, and the boundary 

layer separated between pulses. 

 The combined effect of wakes and vortex generator jets on 

separation control has been studied on the Pack B airfoil by Bloxham 

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distribution is unlimited. ed for public release; 

et al. [41] and on the L1A by Bons et al. [9]. On the L1A, a case was 

considered with Re=50,000, 3% freestream turbulence, and 

dimensionless wake passing frequency F=0.34. The VGJ pulsing 

frequency was equal to the wake passing frequency, and the timing of 

the jets was varied relative to the wakes. Without VGJs the wakes 

caused only partial separation control. With the VGJs injected near 

the pressure minimum on the suction side, good separation control was 

achieved even without wakes. The timing of the jets to the wakes was, 

therefore, unimportant. When the VGJs were injected farther 

downstream, they were only effective when combined with wakes and 

the effectiveness depended on the timing. 

 In the present study, the combined effect of wakes and VGJs is 

considered with Re=25,000 and 50,000 under low (0.6%) and high 

(4%) freestream turbulence conditions. These Reynolds numbers are 

very low, but could still be of interest in small engines operating at 

high altitudes (e.g. in future unmanned vehicles). They are also of 

interest for the present study because they result in a very large 

separation bubble, providing a challenging case for flow control and a 

good case for exploring the response of the boundary layer to VGJs 

and wakes. The wake passing frequency is set to a low value so that 

without flow control the boundary layer only intermittently reattaches 

during wake passing events. Cases with various VGJ frequencies and 

timings relative to the wake passing are documented. Surface pressure 

distributions, total pressure loss profiles, and instantaneous boundary 

layer velocity measurements are used to show how wakes and VGJs 

combine to affect separation. 

EXPERIMENTAL FACILITY AND MEASUREMENTS

 Experiments were conducted in a closed loop wind tunnel with a 

seven blade linear cascade as shown in Fig. 1. A fine screen located 

upstream of the cascade is used to break up the boundary layers which 

form upstream of the test section and to provide uniform inlet 

conditions to the cascade. The freestream turbulence entering the 

cascade was measured with a cross-wire probe positioned just 

upstream of the center blade. The turbulence intensity is 0.8% in the 

streamwise component and 0.5% in the cross stream components. The 

integral length scale of the streamwise component is 0.47Cx. To 

produce high freestream turbulence, the screen is replaced with a 

coarse grid, consisting of a 1.5 mm thick sheet metal plate with 19 mm 

square holes spaced 25.4 mm apart, center to center, in both directions. 

In a plane perpendicular to the inlet flow and 1.7Cx upstream of the 

center blade, the grid produced uniform flow with TI=6.0% in the 

streamwise component and 4.2% in the cross stream components, for 

an overall intensity of 4.9%. The streamwise component was also 

measured at the inlet plane of the cascade in the four center passages, 

where it had decayed to about 4.2%. Downstream of the cascade, the 

local TI is 1.8% across all passages. The local freestream turbulence 

intensity in the passage at the beginning of the adverse pressure 

gradient region is 1.4%. The change in TI through the passage is due 

mainly to the change in the local freestream velocity along with some 

decay of the turbulence. The upstream integral length scale of the 

freestream turbulence is 0.12Cx in the streamwise component and 

0.04Cx in the other components. Further details of the facility and 

inlet flow are in Volino et al. [33]. 

 A tailboard and two flaps, shown in Fig. 1, are needed to produce 

the correct inlet and exit flow angle from the cascade. Their position 

was set to produce periodicity at high Reynolds numbers as discussed 

in Volino [35]. At low Reynolds numbers, when significant separation 

bubbles are present, the periodicity is not as good due to suppression 

of the separation bubble thickness on the blades closest to the 

tailboard. In cases where wakes or other flow control suppress 

separation, periodicity is reestablished. The lack of periodicity in 

Table 1: Cascade parameters 

Axial 

Chord, Cx

[mm] 

True 

Chord 

[mm] 

Pitch, 

Lφ

[mm] 

Span 

[mm] 

Suction 

side, Ls

[mm] 

Inlet 

flow 

angle 

Exit 

flow 

angle 

134 146 136 724 203 35° 60° 

tailboard

flap

flap

screen

B1

B2

B3

B4

B5

B6

B7

pitot

Fig. 1 Schematic of linear cascade with wake generator 

cases with large separation bubbles is considered acceptable since the 

focus of the research is separation control, and not documentation of 

cases with large separation that would be unacceptable in practice. 

This compromise facilitates the study of a larger number of cases by 

obviating the need to adjust the tailboard by trial and error for each 

case. It also provides for better repeatability in the experiments, since 

the position of the tailboard is fixed for all cases. Any changes in 

separation with wakes or VGJs will be larger in practice than 

documented in the experiment, due to the effect of the tailboard in 

suppressing the bubble size in the uncontrolled cases. 

 The wake generator includes a chain near each endwall of the 

cascade that passes 0.54Cx upstream of the leading edges of the 

cascade blades. The chains then pass downstream around blade B7 on 

the inside turn of the cascade and pass well downstream of the cascade 

before returning upstream around blade B1 on the outside turn of the 

cascade. This completes the chain circuit. The magenta line 

surrounding the cascade in Fig. 1 shows the location of the chain. A 

traverse for probe movement is located within the chain circuit 

downstream of the blade row. Each chain is driven by a drive gear 

(large circle in Fig. 1) and also passes around six idler sprockets (small 

circles). One of the idler sprockets is adjustable to maintain tension in 

the chain. The drive gears for the upper and lower chains are on a 

common axle and driven by a single electric motor so both chains 

move in unison. The motor speed is set with a variable frequency 

inverter. The chain links have hollow pins, through which the wake 

generator rods are attached. Each rod consists of a 4 mm diameter 

carbon fiber tube with a steel pin attached at each end. The steel pins 

are inserted through the holes in the upper and lower chain, and then 

secured with small clips. The distance between rods was 272 mm, 

which correspond to 2Lφ, where Lφ is the blade spacing in the cascade. 

The ratio of rod to blade spacing is at the very high end of what might 

be found for vane to rotor blade spacing in an engine. The ratio of 

vane to rotor blade spacing is typically about 1.6 as indicated by 

Bloxham et al. [41], so the high ratio in the present case provides a 

more challenging case for flow control, as shown in Volino [35]. 

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 The ratio of the rod diameter to the axial chord is 0.03, which is 

consistent with the wake generators of Bons et al. [9] and Funazaki et 

al. [16]. The rods are smaller than those of Kaszeta et al. [42] who had 

a diameter to chord ratio of 0.06. The present rods are larger than 

those of Schobeiri et al. [11] and Zhang and Hodson [15] who had rod 

diameter to chord ratios of about 0.01. In the present study, as in Bons 

et al. [9] and Kaszeta et al. [42], the rod wakes are intended to simulate 

the wakes of very highly loaded airfoils under low Reynolds number 

conditions with thick boundary layers and in some cases large 

separation bubbles. A large diameter rod is therefore needed to 

simulate an airfoil wake with a large velocity deficit. The velocity 

deficit and turbulence level in the rod wakes are documented in Volino 

[35], and compared to the wakes of the cascade airfoil. The rod wakes 

were found to be reasonable approximations of airfoil wakes. At the 

cascade inlet, the peak turbulence level in the rod wakes was 14%, and 

the level between wakes was at the background TI in the wind tunnel. 

 The rods were driven at a velocity of 0.73 times the cascade inlet 

velocity, Ui. This gives a flow coefficient, ζ =Uicos(αi)/Urod=1.13, 

where αi is the inlet flow angle. This is at the high end of the expected 

range for an engine. The flow coefficient and rod spacing were chosen 

to be large to provide cases in which the wakes alone would not 

completely eliminate separation. This allows for investigation of the 

interaction of the wakes and VGJs in controlling separation. The 

dimensionless wake passing frequency is F=0.14. The timing of the 

wake generator is recorded with an infrared photo detector, which 

senses the passage of each rod and emits a voltage that is used to 

trigger a function generator which drives the solenoid valves used to 

produce the pulsed VGJs. The signal to the valves is recorded with 

other data, allowing phase averaging of the results. 

 To produce the VGJs, each blade in the cascade has a central cavity 

which extends along the entire span. As explained in Volino et al. 

[36], compressed air is supplied to the cavities from a common 

manifold. The manifold is supplied through two fast response 

solenoid valves (Parker Hannifin 009-0339-900 with General Valve 

Iota One pulse driver) operating in parallel. A single spanwise row of 

holes was drilled into the suction surface of each blade at the inviscid 

pressure minimum location, s/Ls=0.5 (x/Cx=0.62), where s is the 

distance from the leading edge and Ls is the suction surface length. 

The pressure minimum has been shown in the studies listed above to 

be about the optimal location for flow control devices. The holes are 

0.8 mm (0.006Cx) in diameter and drilled at 30° to the surface and 90° 

to the main flow direction. This is the same orientation used in all the 

VGJ studies listed above. The hole spacing is 10.6 diameters, and the 

length to diameter ratio is 12. When the solenoid valves are opened, 

the jet velocity rises quickly for about 0.01 s to a maximum and then 

immediately begins to drop. If the period of the pulse is long enough 

the velocity reaches a steady value, but for the short duration pulses of 

the present study, the 0.01 s rise time compares to valve-open times 

between about 0.01 s and 0.03 s, so there is insufficient time for the jet 

velocity to reach a steady value. When the valves close, the jet 

velocity quickly drops to zero. The maximum jet velocity in each 

pulse is used to define the blowing ratio and is set equal to the nominal 

local freestream velocity at the VGJ holes, for a blowing ratio of 1. 

The jet duty cycle is 10%. The mass flow rate of the jets is 

approximately 0.004% of the main flow mass flow rate. More on the 

characteristics of the pulses is available in Volino et al. [36]. 

 Nine different VGJ timings relative to the wakes were considered, 

as shown in Fig. 2. Timings were chosen to place the jet pulses at 

different times within or between wakes. Case (a) is the baseline case 

with only wakes. Cases (b-d) have a single pulse for each wake 

passing. Timings (e-g) have two pulsed per wake period. To maintain 

Table 2: Velocity profile measurement stations 

Station 1 2 3 4 5 6 

s/Ls 0.53 0.59 0.69 0.78 0.88 0.97 

x/Cx 0.65 0.72 0.80 0.86 0.92 0.97 

Fig. 2 VGJ timings (a) through (i) 

the same overall blowing period, the pulses in cases (e-g) are half as 

long as those in (b-d). Timings (h-i) have three pulses per wake. 

Measurements 

 The center blade, designated B4 in Fig. 1, contains pressure taps 

near the spanwise centerline. Pressure surveys are made using a 

pressure transducer (0-870 Pa range Validyne transducer). Stagnation 

pressure is measured with a pitot tube upstream of the cascade and 

wake generator. The uncertainty in the suction side pressure 

coefficients, Cp, is 0.07. Most of this uncertainty is due to bias error. 

Stochastic error is minimized by averaging pressure transducer 

readings acquired at a 10 kHz sampling rate over a 10 second period. 

 Total pressure losses are documented using a Kiel probe traversed 

across three blade spacings, 0.63Cx downstream of the cascade. A 

traverse is located in the wind tunnel downstream of the cascade to 

move the probe. The traverse causes an acceptably low blockage 

when it is located at least two Cx downstream of the cascade. 

 Velocity profiles on the suction surface were measured at the six 

streamwise stations listed in Table 2. All stations are downstream of 

the inviscid pressure minimum at s/Ls=0.49. Profiles were acquired 

near the spanwise centerline of the airfoil with a hot-wire anemometer 

(AA Lab Systems model AN-1003) and a single sensor hot-film probe 

(TSI model 1201-20). The sensor diameter is 51 μm, and the active 

length is 1.02 mm. At each measurement location, data were acquired 

for 26 seconds at a 20 kHz sampling rate (219 samples). All raw data 

were saved. The high sampling rate provides an essentially continuous 

signal, and the long sampling time results in low uncertainty in both 

statistical and spectral quantities. Data were acquired at 40 wall 

normal locations in each profile, extending from the wall to the freestream, with most points concentrated in the near wall region. The 

probe was positioned as close to tangent to the airfoil surface as 

possible at each station, such that the probe body extended 

downstream of the sensor and the direction of the traverse was within 

5° of normal to the surface. In most cases the closest point to the wall 

in each profile was within about 0.2 mm of the wall, which compares 

to boundary layer thicknesses ranging from 1.1 mm to over 40 mm. 

 Flow direction in a separation bubble cannot be determined with a 

single-sensor hot-wire, but velocity magnitude can be measured and 

was found to be near zero within the bubbles of the present cases when 

the flow was laminar. In cases where the flow became turbulent but 

remained separated, fluctuating velocities caused false high mean 

velocity readings in the separation bubble. With the exception of these 

turbulent separated cases, the uncertainty in the mean velocity is 3-5% 

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except in the very near wall region, where near-wall corrections (Wills 

[43]) were applied to the mean velocity. 

 Velocity was also measured downstream of the cascade along the 

same line used for the total pressure loss measurements. Downstream 

and boundary layer velocity data were both time averaged and 

ensemble averaged based on the phase within the wake passing period. 

Phase averages of mean and fluctuating velocity are shown below at 

24 dimensionless times, t/T, within the wake passing period, where t is 

time and T is the period between wakes. With the wake passing 

frequency of F=0.14 (corresponding to 3 Hz at Re=25,000 and 6 Hz at 

Re=50,000) and a 26 s data acquisition length at each measurement 

location, 78 and 157 wake passing periods are averaged for each 

ensemble for the Re=25,000 and 50,000 cases respectively. 

 Data were acquired at nominal Re=25,000 and 50,000. The 

Reynolds number, as defined above, is based on the suction surface 

length and the nominal cascade exit velocity. The corresponding 

Reynolds numbers based on the cascade inlet velocity and the axial 

chord length are 10,000 and 20,000. For Re=25,000, data were 

acquired for timings (a) and (c-i). For Re=50,000, timings (a-f) were 

used. Data were acquired for cases with both high and low freestream 

turbulence. 

RESULTS 

Re=25,000 

 The Cp and total pressure loss profiles for cases with Re=25,000 

and low TI are shown in Fig. 3. The integrated total pressure losses 

for all cases are shown in Fig. 4. As explained in Volino [31], the 

integrated loss is an average of the loss coefficient across one blade 

spacing centered on the blade B4 wake. The inviscid Cp profile for 

the L1A airfoil is shown in Fig. 3a for comparison. The low peak in 

Cp followed by a plateau in the case without wakes indicates 

separation without reattachment. Wakes alone (case (a)) do not cause 

much change, although there is a slight drop in Cp near the trailing 

edge, indicating possible reattachment for part of the wake passing 

cycle. Little change is observed for cases (c-f). The corresponding 

loss profiles in Fig. 3b show no significant change from the no-wake 

case for cases (a-f) indicating that neither the wakes nor VGJs are 

effective in controlling separation. Case (g), which has two pulses 

evenly spaced between wakes, has slightly more of a drop in Cp near 

the trailing edge, which would suggest more reattachment, but no 

change is observed in the loss profile. Cases (h-i) show better Cp

results with three pulses. In case (h) one of the VGJ disturbances 

coincides with the wake disturbance, while for case (i) all three pulses 

lie between wakes. The loss profiles for cases (h-i) agree with the Cp

results, with a noticeable drop in the loss peaks, better periodicity 

across the cascade, and a shift in the peaks to the right. The shift 

corresponds to an increase in flow turning of about 3°. For the low TI, 

Re=25,000 cases, the variation in the integrated loss, ψint, among cases 

(a)-(i) is about 0.1, which is of the order of the measurement 

uncertainty. Although cases (h) and (i) show some improvement over 

the other cases, the Cp profiles are still far from reattached flow 

behavior, the loss profiles are still very high, and the flow turning is 

still well below the design point. The results suggest only partial or 

intermittent reattachment in cases (h) and (i), in agreement with the 

velocity results shown below. The timing of the VGJ pulses does not 

appear to matter. Cases (g) and (h) both have two pulses between 

wakes, but case (h) shows better results due to an extra pulse, even 

though this pulse occurs during a wake. Cases (h) and (i) both have 

three pulses, and they have similar results, even though one of the 

pulses in case (h) occurs during the wake. Comparison of the present 

results to a high Reynolds number (Re=200,000) case from Volino 

0 0.2 0.4 0.6 0.8 1

0

0.5

1

1.5

2

2.5

s/Ls

Cp

No wake

a

c

d

e

f

g

h

i

inviscid

a)

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4

B5 B6

No wake

a

c

d

e

f

g

h

i

High Re

b)

Fig. 3 Pressure profiles for low TI, Re=25,000 cases: a) Cp, 

b) Total pressure loss 

No wake a b c d e f g h i High Re 0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

ψint

Case

Low TI, Re=25,000

High TI, Re=25,000

Low TI, Re=50,000

High TI, Re=50,000

Fig. 4 Integrated total pressure loss 

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[31] shows that even in the best of the present cases, the loss peaks are 

over twice as high and there is 13° less flow turning than at high Re. 

Comparison to cases with VGJs but without wakes in Volino et al. 

[36] shows that loss peaks at Re=25,000 can be lowered another 10% 

and the flow turning increased another 6° if the pulsing frequency is 

increased to F=0.56. In cases (h-i) the average VGJ frequency is 0.42. 

The average disturbance frequency (wakes plus VGJs) in case (i) is 

0.56. Comparison to cases with wakes but without VGJs in Volino 

[35] show similar improvements over the present results if the wake 

passing frequency is increased. These results suggest that the 

combination of wakes and VGJs is not particularly beneficial. Either 

can result in some separation control if the frequency is sufficiently 

high, but the presence of the other does not necessarily improve the 

results even if their combined frequency is relatively high. 

 Figure 5 shows the Cp and loss results for the high freestream 

turbulence Re=25,000 cases. Comparison of Figs. 4 and 5 shows 

virtually no differences. The integrated loss of Fig. 4 shows the same 

agreement. Boundary layer velocity results show the same similarity 

between the corresponding high and low TI cases. The same similarity 

between high and low TI is observed for all of the Re=50,000 cases. 

Previous results with wakes alone [35] or VGJs alone [36] showed that 

freestream turbulence effects were small compared to wake or VGJ 

effects on boundary layer behavior. The present results show that this 

small effect is reduced to near zero when both wakes and VGJs are 

present. Because the results for both background TI levels are the 

same, only the low TI results will be presented for the remainder of 

this paper. One result that is unexpected in Fig. 5 is the near zero loss 

value between wakes, as some loss should be caused by the decay of 

the high freestream turbulence, as seen in the high Re case. The near 

zero values may be due to bias error within the experimental 

uncertainty of 0.07. 

 Time averaged mean and fluctuating boundary layer velocity 

profiles are shown in Fig. 6 for the six streamwise measurements 

stations of Table 2. Results agree with the pressure data of Fig. 3. 

The separation bubble is thick for the case without wakes or VGJs. 

The addition of wakes (case (a)) reduces the bubble thickness only 

slightly. Cases (c-g) are essentially identical to each other and have a 

slightly thinner separation bubble than case (a). Cases (h-i), with three 

pulses per wake, are nearly identical to each other and show a 

noticeably thinner separation bubble than the other cases. It should be 

noted, however, that the bubble thickness is still quite thick even in 

cases (h-i). This agrees with the high loss and reduced flow turning 

compared to high Re results shown in Fig. 3. Comparison to cases in 

Volino and Ibrahim [38] shows that a reduced bubble thickness is 

possible if the VGJ pulsing frequency is increased to 0.56. 

 Figure 7 shows phase averaged mean velocity for cases (a), (c), and 

(d). The six columns correspond to the six streamwise stations, and 

the rows correspond to a few representative phases in the wake passing 

cycle. When the separation bubble is most distinct, the measured near 

wall velocity is low and nearly constant, but non-zero due to the 

inability of a single sensor hot-wire probe to distinguish direction in a 

reversing flow. When the boundary layer begins to reattach, the 

velocity profile goes more continuously toward zero at the wall. Cases 

(c) and (d) have a single VGJ pulse, and separation control is limited. 

Differences between the three cases in Fig. 7 are slight, but 

discernable. The boundary layer separates at Station 2 in all three 

cases, and the separation bubble grows at Station 3. The profiles for 

the three cases are nearly identical at Station 2. At Stations 4-6, the 

wake is causing some reattachment at t/T=0-0.333, and the profiles for 

the three cases are in close agreement with each other. At t/T=0.5-

0.583, case (d) shows slight signs of reattachment at the downstream 

stations due to a VGJ pulse, while the other cases are more separated. 

0 0.2 0.4 0.6 0.8 1

0

0.5

1

1.5

2

2.5

s/Ls

Cp

No wake

a

c

d

e

f

i

inviscid

a)

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4

B5 B6

No wake

a

c

d

e

f

i

High Re

b)

Fig. 5 Pressure profiles for high TI, Re=25,000 cases: a) Cp, 

b) Total pressure loss 

0 1 0 1 0 1 0 1 0 1 0 1 0

0.1

0.2

U/Ue

y/Ls

St 12 3 4 5 6

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0

0.1

0.2

u′/Ue

y/Ls

No wake a c−f g h i

Fig. 6 Time averaged velocity profiles at six streamwise 

stations for low TI, Re=25,000: top – mean, bottom – rms 

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NASA/CR—2012-217415 145
0

0.1 St. 1 2 3 4 5 6

y/Ls

St. 1 2 3 4 5 6

0.25

t/T=

St. 1 2 3 4 5 6

0

0.1

y/Ls

0.5

0

0.1

y/Ls

0.75

0 1 0 1 0 1 0 1 0 1 0 1 0

0.1

U/Ue

y/Ls

0.917

Fig. 7 Phase averaged mean velocity profiles for Re=25,000 

cases, columns for six streamwise stations, rows for 

phases in pulsing cycle: blue – case (a), green – case (c), 

red – case (d) 

0

0.1

St. 1 2 3 4 5 6

y/Ls

0.25

t/T=

St. 1 2 3 4 5 6

0.25

0 1 0 1 0 1 0 1 0 1 0 1 0

0.1

U/Ue

y/Ls

0.833

Fig. 8 Phase averaged mean velocity profiles for Re=25,000 

cases: blue – case (a), green – case (g), red – case (h), 

black – case (i) 

At t/T=0.667-0.75 all cases show a thick separation bubble. At 

t/T=0.833, case (c) shows reattachment at Station 3 due to a VGJ 

pulse, and this continues to Station 4 at t/T=0.917 and Station 5 at 

t/T=1. The profiles of Fig. 7 show that the VGJ pulses clearly do have 

some effect, albeit small at this Re, and the appearance of the effect 

depends on the jet timing, although overall separation control is very 

limited, as shown by the time averaged profiles of Fig. 6. 

 Figure 8 shows the effect of multiple VGJ pulses on the phase 

averaged velocity. Case (a), shown for reference, has only wakes. 

Case (g) has two pulses between the wakes. Case (h) has three pulses, 

with two occurring at about the same timings as in case (g) and the 

third coinciding with the wake. Case (i) has three pulses between 

wakes. Comparison to Fig. 7 shows that separation control is better 

with multiple jet pulses, in agreement with the time averaged profiles 

of Fig. 6. Cases (a) and (g) are in close agreement for t/T=0-0.333, 

which corresponds to the wake passing. At later times, case (g) shows 

a somewhat thinner separation bubble than case (a) as the VGJs help 

to control separation. With some minor exceptions, cases (h) and (i) 

are in good agreement with each other for the full cycle. The addition 

of a third VGJ pulse reduces separation from case (g) at all phases, 

both during and between wakes. Whether the third pulse occurs 

during or between wakes does not seem to matter. With three pulses 

the separation bubble is present at Stations 2-4, but by the downstream 

stations the VGJ pulses occur frequently enough to prevent a distinct 

bubble from forming. The strong adverse pressure gradient still results 

in a thick boundary layer in cases (h-i), and at some phases the 

boundary layer is on the verge of separating. 

 The time averaged boundary layer thickness, δ99.5, is shown in Fig. 

9 for the Re=25,000 cases, and these values are used in Fig. 10, which 

shows the phase averaged separation bubble thickness as a fraction of 

the local, time averaged δ99.5. The bubble thickness is estimated as the 

distance from the wall to the farthest point in the shear layer with 

∂u y / ∂ <0. The bubble thickness as a fraction of the boundary layer 

thickness is used to show the local extent of separation. One could 

also use the shape factor, but the uncertainty in the displacement and 

momentum thicknesses resulting from the inability of the hot wire to 

accurately measure velocity within the separation bubble makes the 

shape factor a somewhat less reliable quantity than the bubble 

thickness used here. The dimensional bubble thickness at any given 

time and location can be determined using the data in Figs. 9 and 10 

together. In each time-space plot of Fig. 10, the data are repeated for 

two cycles to show the periodicity. The solid and dashed white lines 

indicate the leading and trailing edges of the wake affected regions. 

The magenta lines bound the VGJ affected regions. In all cases, the 

suppression of the bubble by both the wake and VGJs is clear. As 

shown in previous studies ([4], [38], [37]), the initial transient at the 

start of a VGJ pulse is most effective for flow control. For the wakes, 

it appears that the separation control is effective for the full duration of 

the wake. Hence, the separation bubble is suppressed for about three 

times as long by each wake than by each VGJ pulse. Cases (g) and (h) 

appear very similar in Fig. 10, which would suggest that the extra 

pulse within the wake of case (h) is ineffective. The normalizing 

quantity, δ99.5 is lower for case (h), however, so the extra pulse does 

help, as shown in Figs. 3, 6 and 8. 

 The effects of the boundary layer behavior on the downstream flow 

are shown in Figs. 11 and 12, which show the phase averaged mean 

and fluctuating velocity 0.63Cx downstream of the blade row. The 

phase averaged mean velocity is the average of all the velocity data at 

a particular location and phase. The fluctuating velocity is the rms of 

the difference between the instantaneous velocity and the local phase 

averaged mean at each location and phase. Cases (a) and (i) are shown 

as examples. The contours are normalized by the exit velocity Ue. 

Since the rod spacing is 2Lφ, the flow in alternating passages is in 

phase, with the passages between a half cycle out of phase. In Figs. 

11a and 12a, there are vertical strips of low mean velocity at φ/Lφ = 

-1.5, -0.5, 0.5 and 1.5 which result from the velocity deficit in the 

airfoil wakes. Strips of high fluctuating velocity are present at these 

same locations in Figs. 11b and 12b. These positions correspond to 

the loss peaks in Fig. 3b. The dimensionless mean velocity between 

the airfoil wakes cycles between a low of about 0.85 and a high of 

about 1.1 showing the velocity deficit in the rod wakes and the 

acceleration between wakes. The rod wakes proceed at an angle in the 

figure, rising from left to right, as they move forward in time and 

transit across the cascade. The highest turbulence peaks occur where 

the rod wakes interact with the separation bubble and airfoil wakes. 

Comparing Figs. 11 and 12, the amount of mean velocity variation and 

turbulence in case (i) is less than in case (a) as the VGJs reduce the 

amount of separation and reattachment in the boundary layer. The 

separation control, as shown above, is not complete, however, so the 

7

NASA/CR—2012-217415 146
0.5 0.6 0.7 0.8 0.9 1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

0.2

0.22

s/Ls

δ99.5/Ls

a

c

d

e

f

g

h

i

Fig. 9 Time averaged boundary layer thickness, δ99.5, for 

Re=25,000 cases 

results for case (a) with wakes only and case (i) with the best control 

of the cases considered at Re=25,000, are not drastically different. 

Larger differences were observed in Volino [35] in cases where higher 

wake passing frequencies resulted in better separation control. 

Re=50,000 

 When the Reynolds number is increased to 50,000, the boundary 

layer is less prone to separation and more easily controlled. As noted 

above, full data sets were acquired at Re=50,000 for both high and low 

background freestream turbulence, but since the high and low TI

results are nearly indistinguishable, only the low TI results are 

presented below. The Cp and total pressure loss profiles are shown in 

Fig. 13. Without wakes or VGJs, the boundary layer separates and 

does not reattach. With wakes alone (case (a)) the Cp profile shows a 

large drop near the trailing edge, indicating reattachment, but the peak 

Cp value is below the value in the better controlled cases, indicating 

that reattachment is likely not complete for the full cycle. Cases (b), 

which has a single VGJ pulse occurring coincident with the wake, 

matches case (a), indicating that the single pulse provides no benefit 

when it overlaps the wake. In cases (c) and (d), which both have a 

single VGJ pulse between wakes, the separation control is better. The 

Cp profiles suggest a bubble is present between s/Ls=0.6 and 0.8, but 

the boundary layer is attached farther downstream. With two pulses 

per wake in cases (e) and (f), results are similar, but reattachment 

appears to move slightly upstream. The loss profiles of Fig. 12b are 

consistent with the Cp results, particularly for the center blade, B4. 

Cases (a) and (b) are similar to each other and have loss peaks about 

30% lower than in the case without wakes. This is consistent with the 

integrated results shown in Fig. 4. The peaks are also shifted to the 

right of the no-wake case, indicating an increase in flow turning of 

about 3°. For cases (c-f) there is an additional drop in the loss peak 

and a further shift to the right, indicating about 8° more flow turning 

than the no-wake case. Even in the best case the loss peak is much 

larger and indicates about 5° less flow turning than in the high 

Reynolds number comparison case. For cases (e) and (f), which have 

the best separation control, the wakes of blades B4 and B5 appear 

similar. For the other cases, the control is partial and as noted above 

the tailboard has an effect of further suppressing separation on the 

closer blades, resulting in poorer periodicity. This effect may be more 

apparent at Re=50,000 than in the Re=25,000 case of Figs. 3 and 5 

a

t/T

0.5

1

1.5

2

c

t/T

d

0.5

1

1.5

2

e

t/T

f

0.5

1

1.5

2

g

s/Ls

h

t/T

0.6 0.7 0.8 0.9

0.5

1

1.5

2

s/Ls

i

0.6 0.7 0.8 0.9

0 0.2 0.4

Fig. 10 Time space plots of phase averaged separation 

bubble thickness normalized on local time averaged δ99.5 for 

Re=25,000 cases: solid white line – leading edge of wake, 

dashed white line – trailing edge of wake, magenta lines – 

extent of VGJ affected region 

since Re=50,000 is a borderline case with respect to separation, and 

the boundary layer can go from fully separated to nearly fully attached 

with relatively small changes in conditions. This appears to make the 

boundary layer more sensitive to the influence of the tailboard and 

leads to differences between passages. At Re=25,000, separation 

control was only partial even in the best cases. The tailboard may act 

to reduce the bubble thickness more in some passages than others, but 

it is less likely to produce reattachment. Hence there tends to be more 

variability between passages at Re=50,000 than 25,000. 

8

NASA/CR—2012-217415 147
φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

a)

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

b) 

Fig. 11 Time space plot of phase averaged velocity 0.63Cx

downstream of cascade for Re=25,000 case (a): a) U/Ue, b) 

u /Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

a)

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

b)

Fig. 12 Time space plot of phase averaged velocity 0.63Cx

downstream of cascade for Re=25,000 case (i): a) U/Ue, b) 

u /Ue

0 0.2 0.4 0.6 0.8 1

0

0.5

1

1.5

2

2.5

s/Ls

Cp

No wake

a

b

c

d

e

f

inviscid

a)

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4

B5 B6

No wake

a

b

c

d

e

f

High Re

b)

Fig. 13 Pressure profiles for low TI, Re=50,000 cases: a) 

Cp, b) Total pressure loss 

0 1 0 1 0 1 0 1 0 1 0 1 0

0.1

0.2

U/Ue

y/Ls

St 12 3 4 5 6

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0

0.1

0.2

u′/Ue

y/Ls

No wake a b c−d e f

Fig. 14 Time averaged velocity profiles at six streamwise 

stations for low TI, Re=50,000: top – mean, bottom – rms 

9

NASA/CR—2012-217415 148
10 

This material is declared a work of the U.S. Government and is

 not subject to copyright protection in the United States. Approv

distribution is unlimited. ed for public release; 

 Time averaged velocity profiles for the Re=50,000 cases are shown 

in Fig. 14. Without wakes there is a large separation bubble, in 

agreement with the pressure results of Fig. 13. With wakes alone in 

case (a), the separation bubble is slightly thinner. In cases (b-d), 

which all have one VGJ pulse per wake, the separation bubble is much 

thinner. At Station 6, there is no clear separation bubble, although the 

boundary layer appears to be on the verge of separating. In the mean 

profiles, cases (b-d) are virtually indistinguishable from each other. In 

the fluctuating profiles, the peak in case (b) is slightly farther from the 

wall than in cases (c-d). This result agrees with the trend in Fig. 13, 

which shows that separation control is slightly less effective when the 

pulse coincides with the wake, but the differences associated with the 

timing are small and are not apparent in all measured quantities. In 

cases (e-f), which have two pulses per wake, the separation bubble is 

slightly thinner than in cases (b-d), and the peak in the fluctuating 

velocity is smaller and closer to the wall. 

 The boundary layer thicknesses, determined from the profiles of 

Fig. 14, are shown in Fig. 15. Figure 16 shows the phase averaged 

separation bubble thickness as a fraction of δ99.5. The wake and VGJ 

pulses both suppress separation at all locations. Between wakes and 

pulses the boundary layer separates at s/Ls 0.7 in all cases. Without 

VGJs (case (a)) this separation persists to the trailing edge. With 

VGJs, regardless of timing, the boundary layer reattaches at all phases. 

 Figures 17 and 18 show the phase averaged mean and fluctuating 

velocity in the airfoil wakes for cases (a) and (e). As in Figs. 11 and 

12, the velocity deficits and turbulence in the airfoil wakes are clear at 

φ/Lφ = -1.5, -0.5, 0.5 and 1.5. The rod wakes appear between the 

airfoil wakes as areas of slightly elevated fluctuating velocity in Figs. 

16b and 17b at φ/Lφ = -1, and 1 when t/T=0.15; and at φ/Lφ = 0 when 

t/T=0.65. As in Figs. 11 and 12, the fluctuating velocity is highest 

when the rod wakes interact with the airfoil wakes. The variation in 

the mean velocity and the turbulence are lower for case (e) than case 

(a) because the VGJs reduce the growth of the separation bubble. The 

reduction in wake strength and turbulence is in agreement with the 

thinner boundary layer and separation bubble shown in Figs. 14-16, 

and the reduced losses of Fig. 13b. Wake measurements for the other 

Re=50,000 cases show the same trend. 

 The results at both Re=25,000 and 50,000 show that when acting 

together, wakes and VGJs suppress separation. This result is expected 

since wakes and VGJs were both previously shown to suppress 

separation when acting alone. Somewhat surprisingly, the boundary 

layer was not very sensitive to the timing of the VGJs with respect to 

the wakes. It had been expected that VGJs timed to pulse between 

wakes would help suppress the growth of the separation bubble 

between the wakes, while VGJs timed to coincide with the wakes 

might be wasted since the wake would already be acting to suppress 

separation at the instant when the jets were pulsed. Cases (b-d) at 

Re=50,000 indicate that timing the VGJs to avoid the wakes may have 

some benefit, but the influence of timing is small. A pulse lying 

completely within a wake (case (b)) still helps suppress separation and 

reduce losses, and is nearly as effective as a pulse between wakes 

(cases (c) and (d)). For the Re=25,000 cases, two pulses between 

wakes (case (g)) were not enough to significantly reduce separation 

and losses, but when a third pulse was added within the wake (case 

(h)), separation was reduced. In fact, case (h) was just as effective as 

case (i), which had all three pulses between wakes. 

CONCLUSIONS 

 The combined effects of unsteady wakes and vortex generator jets 

on the flow over the very high lift L1A airfoil were studied 

experimentally under low and high freestream turbulence conditions. 

0.5 0.6 0.7 0.8 0.9 1

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

s/Ls

δ99.5/Ls

a

b

c

d

e

f

Fig. 15 Time averaged boundary layer thickness, δ99.5, for 

Re=50,000 cases 

t/T

a

0.5

1

1.5

2

b

t/T

c

0.5

1

1.5

2

d

s/Ls

t/T

e

0.6 0.7 0.8 0.9

0.5

1

1.5

2

s/Ls

f

0.6 0.7 0.8 0.9

Fig. 16 Time space plots of phase averaged separation 

bubble thickness normalized on local time averaged δ99.5 for 

Re=50,000 cases: solid white line – leading edge of wake, 

dashed white line – trailing edge of wake, magenta lines – 

extent of VGJ affected region (same color scale as Fig. 9) 

10 NASA/CR—2012-217415 149
φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

a)

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

b) 

Fig. 17 Time space plot of phase averaged velocity 0.63Cx

downstream of cascade for Re=50,000 case (a): a) U/Ue, b) 

u /Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

a)

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

b)

Fig. 18 Time space plot of phase averaged velocity 0.63Cx

downstream of cascade for Re=50,000 case (e): a) U/Ue, b) 

u /Ue

Reynolds numbers based on suction surface length and nominal exit 

velocity of 25,000 and 50,000 were considered. The effect of the 

background freestream turbulence between wakes was negligible in 

the presence of larger wake and VGJ disturbances. Results for cases 

with TI=0.6% and 4% were nearly identical. The wake passing 

frequency considered in the present study was sufficiently low that the 

wakes caused only intermittent reattachment with a large, unclosed 

separation bubble appearing between wakes. Vortex generator jets 

were able to help reduce this separation if their pulsing frequency was 

sufficiently high. The timing of the jets with respect to the wakes was 

not particularly important. For the cases considered, the beneficial 

effects of the wakes and VGJs did not appear to be additive. The jet 

pulsing frequency required to fully control separation was about the 

same as needed in cases without wakes. 

ACKNOWLEDGMENTS

 This work was sponsored by the National Aeronautics and Space 

Administration under grant NNC07IA10I. The grant monitors were 

Drs. Anthony Strazisar and James Heidmann of the NASA Glenn 

Research Center. The support of the United States Naval Academy 

Technical Support Department Shop and Fluids Laboratory is greatly 

appreciated. 

 

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Loaded Low-Pressure Turbine Airfoil,” ASME Journal of 

Turbomachinery, 132, 011007.

[32] Ibrahim, M., Kartuzova, O., and Volino, R.J., 2008, 

“Experimental and Computational Investigations of Separation 

and Transition on a Highly Loaded Low Pressure Turbine Airfoil: 

Part 1 – Low Freestream Turbulence Intensity,” ASME Paper 

IMECE2008-68879. 

[33] Volino, R.J, Kartuzova, O., and Ibrahim, M., 2008, 

“Experimental and Computational Investigations of Separation 

and Transition on a Highly Loaded Low Pressure Turbine Airfoil: 

Part 2 – High Freestream Turbulence Intensity,” ASME Paper 

IMECE2008-68776.

[34] Volino, R.J., 2002, “Separated Flow Transition Under Simulated 

Low-Pressure Turbine Airfoil Conditions: Part 1 – Mean Flow 

and Turbulence Statistics,” ASME Journal of Turbomachinery, 

124, pp. 645-655. 

[35] Volino, R.J., 2010, “Effect of Unsteady Wakes on Boundary 

Layer Separation on a Very High Lift Low Pressure Turbine 

Airfoil,” ASME Paper GT2010-23573. 

[36] Volino, R.J., Kartuzova, O., and Ibrahim, M.B., 2009, 

“Experimental and Computational Investigations of Low-Pressure 

Turbine Separation Control Using Vortex Generator Jets,” ASME 

Paper GT2009-59983. 

[37] Volino, R.J., Kartuzova, O., and Ibrahim, M.B., 2010, 

“Separation Control on a Very High Lift Low Pressure Turbine 

Airfoil Using Pulsed Vortex Generator Jets,” ASME Paper 

GT2010-23567. 

[38] Volino, R.J., and Ibrahim, M.B., 2010, “Separation Control on 

High Lift Low Pressure Turbine Airfoils Using Pulsed Jet Vortex 

Generator Jets,” Proceedings of the ASME-ATI-UIT 2010 

Conference on Thermal and Environmental Issues in Energy 

Systems, 2010. 

[39] Ibrahim, M., Kartuzova, O., Doucet, D., and Volino, R.J., 2010, 

“LES Flow Control Simulations for Highly Loaded Low Pressure 

Turbine Airfoil (L1A) Using Pulsed Vortex Generator Jets,” 

ASME paper GT2010-23015. 

[40] Ibrahim, M.B., Kartuzova, O., and Volino, R.J., 2009, “LES and 

URANS Computational Investigations of LPT Blade (L1A) 

Separation Control using Vortex Generator Jets,” Proceedings of 

Turbulence, Heat and Mass Transfer 6, 2009. 

[41] Bloxham, M., Reimann, D., Crapo, K., Pluim, J., and Bons, J.P., 

2007, “Synchronizing Separation Flow Control with Unsteady 

Wakes in a Low-Pressure Turbine Cascade,” ASME Paper 

GT2007-27529. 

[42] Kaszeta, R.W., Simon, T.W., Jiang, N., and Ottaviani, F., 2005, 

“Influence of Wake Passing Frequency and Elevated Turbulence 

Intensity on Transition,” AIAA Journal of Thermophysics and 

Heat Transfer, 19, 137-147. 

[43] Wills, J.A.B., 1962, “The Correction of Hot-Wire Readings for 

Proximity to a Solid Boundary,” Journal of Fluid Mechanics, 12, 

pp. 65-92. 

12 NASA/CR—2012-217415 151
“The United States Government retains, and by accepting the article for publication, the publisher 

acknowledges that the United States Government retains, a non-exclusive, paid-up, irrevocable, 

worldwide license to publish or reproduce the published form of this work, or allow others to do so, for 

United States Government purposes." 

CFD SIMULATIONS OF UNSTEADY WAKES ON A HIGHLY LOADED LOW 

PRESSURE TURBINE AIRFOIL (L1A)

Mounir B. Ibrahim and Samuel Vinci

Mechanical Engineering Department

Cleveland State University

Cleveland, Ohio 44115-2425

m.ibrahim@csuohio.edu

Olga Kartuzova

National Center for Space 

Exploration Research 

NASA Glenn Research Center

Cleveland, Ohio 44135 

olga.kartuzova@nasa.gov

Ralph J. Volino

Mechanical Engineering Department 

United States Naval Academy

Annapolis, Maryland 21402-5042 

volino@usna.edu

ABSTRACT

A study of a very high lift, low-pressure turbine airfoil in 

the presence of unsteady wakes was performed 

computationally and compared against experimental results. 

The experiments were conducted in a low speed wind tunnel 

under high (4.9%) and then low (0.6%) freestream turbulence 

intensity conditions with a flow coefficient (of 0.7. The 

experiments were done on a linear cascade with wakes that 

were produced from moving rods upstream of the cascade

with the rod to blade spacing varied from 1 to 1.6 to 2. In the 

present study two different Reynolds numbers (25,000 and 

50,000, based on the suction surface length and the nominal 

exit velocity from the cascade) were considered. 

The experimental and computational data have shown that 

in cases without wakes, the boundary layer separated and did 

not reattach. The CFD was performed with Large Eddy 

Simulation (LES) and Unsteady Reynolds-Averaged NavierStokes (URANS), Transition-SST, utilizing the finite-volume 

code ANSYS FLUENT under the same freestream turbulence 

and Reynolds number conditions as the experiment but only at 

a rod to blade spacing of 1.

With wakes, separation was largely suppressed, 

particularly if the wake passing frequency was sufficiently 

high. Similar effect was predicted by 3D CFD simulations. 

Computational results for the pressure coefficients and 

velocity profiles were in a reasonable agreement with 

experimental ones for all cases examined. The 2D CFD 

efforts failed to capture the three dimensionality effects of the 

wake and thus were less consistent with the experimental data. 

As a further computational study, cases were run to 

simulate higher wake passing frequencies which were not run 

experimentally. The results of these computational cases 

showed that an initial 25% increase from the experimental 

dimensionless wake passing frequency of F=0.45 greatly 

reduced the size of the separation bubble, nearly completely 

suppressing it, however an additional 33% increase on top of 

this did not prove to have much of an effect.

NOMENCLATURE

Cp , pressure coefficient

Cx axial chord length

Cf w/(1/2U), skin friction coefficient 

F fLj-te/Uave, dimensionless wake passing frequency

f frequency

HFSTI High Freestream Turbulence Intensity (4.9%)

Lj-te length of an adverse pressure gradient region on the

suction surface

LFSTI Low Freestream Turbulence Intensity (0.6%)

Ls suction surface length

L blade spacing (pitch)

P pressure

PS upstream static pressure

PT upstream stagnation pressure

PTe downstream stagnation pressure

Re UeLs/, exit Reynolds number

s streamwise coordinate, distance from leading edge

T period of wake passing cycle

t time

w (dU/dy)y=0, wall shear stress

  2

T

 U/PP2 e

Proceedings of ASME Turbo Expo 2012 

GT2012 

June 11-15, 2012, Copenhagen, Denmark 

GT2012-6

NASA/CR—2012-217415 1521 Copyright © 2012 by ASME This work is in part a work of the U.S. Government. ASME disclaims all interest in the U.S. Government’s contributions.
U local mean velocity

Uave average freestream velocity in adverse pressure 

gradient region, based on inviscid solution

Uiinlet freestream velocity

Ue nominal exit freestream velocity, based on inviscid

solution

u′ rms fluctuating streamwise velocity

x axial distance from leading edge

y wall normal distance from suction surface

 coordinate along blade spacing, normal to axial 

chord

 kinematic viscosity

 density

 (PT-PTe)/(PT-PS), total pressure loss coefficient

 Uicos(i)/Urod = Uaxial/Urod, flow coefficient

INTRODUCTION

The desire to reduce fuel costs and improve engine 

performance in the gas turbine industry has led many to 

conduct experimental and computational fluid dynamics 

(CFD) research on the effects of flow over the low-pressure 

turbine (LPT) blades. This research has shown that separation 

of the flow can occur on the suction surface of these blades 

due to the presence of adverse pressure gradients. Separation 

results in partial loss of lift and higher aerodynamic losses at 

high altitude cruise conditions and becomes increasingly more 

severe as the aft loading of the airfoils increases. 

Early experimental and computational studies of the flow 

over these high lift airfoils were performed by examining the 

airfoils by themselves in a wind tunnel with varying levels of 

freestream turbulence intensities. However, it has become 

evident that the presence of the forward stages of stator vanes 

in front of the rotor airfoil blades creates wake shedding 

vortices, which have significant effect on the boundary layer 

of these rotor blades. The velocity deficit in these wakes 

causes flow outside of the wake to accelerate and impinge on 

the suction surface of blades creating a so called “negative jet” 

Hodson and Howell [1]. The study of this wake and the wake 

induced transition associated with it has started to become the 

topic of study in a number of experiments; however this type 

of flow has yet to be extensively modeled computationally due 

to the difficulties associated with creating these models.

The creation of the impinging wake on a blade has been 

studied by Pluim et al. [2] and it was shown that a circular rod 

provides a sufficient representation of the form of the true 

wake that comes off the upstream vane. Thus experimental 

data are collected by passing a line of circular rods upstream 

of a cascade of airfoils with data collected by hot wire 

anemometers over the suction side of these airfoils. Many 

experiments have been run analyzing the wake/blade 

interaction effects; see Bons et al. [3] and Pluim et al. [2]. 

Present computational study is based on the experiment of 

Volino [4]. Numerical studies of the wake and its subsequent 

interaction with an airfoil was studied by Sarkar [5]. Sakar in 

his study applied a wake that was generated by a circular rod. 

The results of the wake were then interpolated into a domain 

containing the airfoil. The analysis of a circular wake 

generating rod and its effect on a downstream airfoil was 

examined for a Reynolds number of 78,000 (based on the axial 

chord and the inlet velocity). In this study airfoil T106 (see 

Stieger et al. [6]) was considered with flow coefficient (of 

0.83 and reduced blade passing frequency of 0.68. The wake 

was originally generated in both 2D and 3D simulations using 

URANS and LES turbulence models respectively. These 

results were then interpolated into a 3D airfoil geometry and 

the effects of the wake/airfoil interaction were studied with 

LES. It was shown that the wakes produced with LES and 

interpolated into the airfoil domain produced smaller scale 

eddies and thus produced more accurate predictions of 

separated flow than the 2D wakes. 

Suzen and Huang [7] numerically studied unsteady 

wake/blade interactions in LPT Pack-B flows using an 

intermittency transport equation. They followed experiments 

of Kaszeta et al. [8, 9] and Stieger [10]. In that study, 

predictions of the flow with the intermittency transport model 

were in a good agreement with experimental data for the 

pressure coefficient. Differences in velocity predictions in the 

separated region were attributed to the inability of hot wire 

anemometry, used in the experiment, to measure negative 

velocities. The observation was made that high wake passing 

frequency resulted in suppression of the separation zone due to 

higher turbulence levels generated in the wake.

Rodi [11] conducted DNS and LES studies of the flow 

past turbine blades with incoming wakes. Reynolds number 

based on axial chord and inlet velocity was 51,800 in the low 

Re case and 148,000 in the high Re case. In both cases DNS 

and LES showed similar results for the pressure coefficient, 

which were in reasonable agreement with experimental data. 

Some disagreement in pressure coefficient near the leading 

edge on the suction side was attributed to the difference in the 

inlet flow angle and compressibility effects of the flow, which 

was modeled as incompressible. In the high Re case, where 

the flow was attached, LES predicted transition a little later 

compared to DNS and therefore it was concluded that LES is 

not a good tool for predicting this type of flow. However, it is 

necessary to mention that DNS required 10 times longer to 

compute, compared to LES and it cost the author several 

months of calculations on a supercomputer.

Lardeau [12] applied the URANS approach to simulate a 

moving wake generating rod and airfoil blade geometry. He

showed that the URANS models were very sensitive to the 

grid size and the time step size of the calculation. The results 

2

NASA/CR—2012-217415 153
showed that too coarse of a grid or time step size would result 

in the vortex shedding not developing due to the small length 

scale of the circular wake generating rod as compared to the 

blade.

In this paper a computational model was developed and 

applied to simulate effect of the moving upstream rod on the 

flow transition and separation over highly loaded LPT airfoil 

L1A. URANS and LES approaches to turbulence modeling 

were compared. The results were compared with experimental 

data of Volino [4]. CFD provided further insight to the effect 

of the upstream wake on the flow over the airfoil. Additional 

cases were run to analyze effects of rod speed on the boundary 

layer of the airfoil suction surface.

EXPERIMENTAL FACILITY AND MEASUREMENTS

Details of experimental setup are presented in Volino [4] 

and summarized below. 

Experiments were conducted in a closed loop wind tunnel 

with a seven blade linear cascade as shown in Fig. 1 with the 

cascade parameters shown in Table 1. The streamwise 

component of the turbulence intensity was 0.6% with the 

integral length scale of 6.3 cm. A tailboard, shown on Fig. 1, 

was used to ensure the correct flow angle from the cascade 

with periodicity at high Reynolds numbers. At low Reynolds 

numbers the periodicity was not as good because large 

separation was present and due to suppression of the 

separation on the blades closest to the tailboard. 

The wake generator includes a chain near each endwall of 

the cascade that passes 0.54Cx upstream of the leading edges 

of the cascade blades. The chain links have hollow pins, 

through which the wake generator rods are attached with a 

diameter of 4 mm. Tests were run with average distances 

between rods of 136 mm, 221 mm, and 272 mm, which 

correspond to 1Lφ, 1.6Lφ and 2Lφ, where Lφ is the blade 

spacing in the cascade. These ratios of rod to blade spacing are 

in the range expected for vane to rotor blade spacing in an 

engine. 

For most tests, the rods were driven at a velocity of 1.18 

times the cascade inlet velocity, Ui. This gives a flow 

coefficient, ζ =Uicos( αi)/Urod=0.7, where αiis the inlet flow 

angle. This is also in the expected range for an engine. 

Table 1: Cascade parameters

NUMERICAL METHODS

The numerical simulations were conducted utilizing the 

finite-volume code ANSYS FLUENT [13]. Cases were run at 

different Reynolds numbers, wake passing frequencies, and 

free stream turbulence intensity levels. Table 2 shows a 

summary of all CFD cases conducted in this study. 

The computational domain was based on earlier work (see 

Ibrahim et al. [14]). However the grid was sub-divided into 

three cell-zone domains, a stationary inlet zone, a moving 

(sliding, periodically repeating) translational zone containing 

the circular wake generating rod in it, and a stationary zone 

with the airfoil and rest of the downstream domain in it. The 

mesh generated around the rod was refined until it showed a 

wake that was produced and did not dissipate as it propagated 

downstream to the cascade.

Figure 1. Cascade Passage

The interaction between the interface zones was set up so 

the conditions periodically repeated themselves allowing for 

the wake to carry between zones even as the interfaces became 

unaligned. The inlet and outlet conditions varied based on 

Reynolds number and the freestream turbulence intensity of 

the flow and the boundaries on each face of the grid were set 

up as periodic in order to simulate a full blade and a full 

cascade. For the two dimensional cases this assumed no 

variation in the z direction, which proved to cause some

modeling issues. 

For the two dimensional cases, a URANS calculation 

utilizing the Transition-SST turbulence model of Menter et al. 

[15] was used. The three dimensional computational domain 

was used for both the URANS calculations and LES

calculations. The number of passages used was varied for the

Axial 

Chord, Cx

[mm]

True 

Chord 

[mm]

Pitch,

L

[mm]

Span

[mm]

Suction 

side, Ls

[mm]

Inlet 

flow 

angle

Exit 

flow 

angle

134 146 136 724 203 35° 60°

3

NASA/CR—2012-217415 154
Table 2: CFD cases investigated

Experimental Data available from Volino [4].

two dimensional and three dimensional domains and will be 

explained further below (see Table 1. for cascade parameters). 

A uniform velocity inflow condition is specified 1.9 Cx

upstream of the blade leading edge in the flow direction. The 

inlet flow angle is set to 33obased on an inviscid calculation 

of the full cascade used in the experiment (see Ibrahim et al. 

[16]). This angle agrees with the experimentally measured 

inlet angle to within the experiment uncertainty. The exit 

boundary is located 3.8 Cx downstream of the trailing edges in 

the flow direction. The boundary conditions on the sides of the 

passage are periodic. 

Convergence was established when: 1) residuals reduced 

to a value 10-5, 2) no change was observed in any field results, 

and 3) the mass imbalance was less than 0.01 %. 

For the 2D cases, a two-channel domain was designed, 

having an entire airfoil in the middle, with two rods spaced 

equally one blade spacing apart. The details of this grid are 

given in Table 3. In the 3D simulations, with rod to blade 

spacing of 1, a single airfoil with half channel spacing above 

and below was used. Periodic boundary conditions were 

applied to simulate the entire cascade and in order to save on 

computational memory and time. Adding multiple rods to a 

single domain (in the 2D cases) did not increase the mesh size 

significantly nor computational time, therefore two blade 

passages with two rods were used, however for the 3D domain 

one full blade passage (half a blade passage above and below 

an airfoil) was simulated to keep the mesh size at a minimum. 

Table 2 summarizes cases examined in this paper, with the 

same cases also studied experimentally by Volino [4] 

indicated. A combination of variation in Re (25,000 and 

50,000), free stream turbulence intensities (LFSTI and HFSTI) 

and turbulence modeling approach (URANS with TransitionSST model and LES) were considered to match the 

experimental data. Additional cases with increased rod 

speeds, and thus increased dimensionless wake passing 

frequency values (F) were studied. 

TURBULENCE MODELING 

LES with Dynamic Kinetic Energy Subgrid-Scale model

The governing equations employed for LES are obtained 

by filtering the time-dependent Navier-Stokes equations. The 

filtering process effectively filters out eddies whose scales are 

smaller than the grid spacing used in the computations. The 

resulting equations thus govern the dynamics of large eddies. 

The subgrid-scale stresses resulting from the filtering 

operation are unknown, and require modeling. The subgridscale turbulence models in Fluent employ the Boussinesq

hypothesis as in the RANS models.

The dynamic subgrid-scale kinetic energy model, used in 

the present study, is based on the model proposed by Kim and 

Menon [17]. In this model a separate transport equation is 

solved for subgrid-scale kinetic energy. The model constants 

are determined dynamically. The details of the implementation 

of this model in ANSYS FLUENT and its validation are given 

by Kim [18].

Transition-SST (4 equation) model of Menter et al. [15]

A correlation-based transition model was proposed by 

Menter et al. [15]. This model is based on two transport 

equations. The intermittency transport equation is used to 

trigger the transition onset. The transport equation for the 

transition momentum thickness Reynolds number (Ret) is 

used to capture non-local effects of freestream turbulence 

intensity and pressure gradient at the boundary layer edge. 

Outside the boundary layer the transport variable was forced 

to follow the value of Ret given by correlations. Those two 

equations were coupled with the shear stress transport 

turbulence (SST) model. This model is available in ANSYS 

FLUENT Version-12 [13] as Transition-SST (4 equation) 

turbulence model.

RESULTS AND DISCUSSION

Code Validation

Four different grids were compared in this study as shown 

in Table 3. Velocity magnitude contours were analyzed for the 

2D, Re = 50,000 case comparing Grids#1 and #2. The CFD 

data from Grid#2 & #3 did not vary much and Grid#3 only 

increased computation time. Based on the above Grid#2 was 

selected for 2D simulations of the present study. Grid#4 was 

generated for 3D simulations of the present study. This grid is 

similar to the Grid#2 and expanded in the spanwise direction.

Grid#4 is suitable for LES since grid resolution is within 

recommended values for this type of simulation (in the nearwall region recommended values are: y+~ 2; x+ ~ 50-150; 

Case Re

Free Stream 

Turbulence 

Intensity

F

Turbulence

Model

1 25,000 LFSTI* 0.45 Trans-SST 

2 50,000 0.45 (2D)

3

25,000 LFSTI* 0.45

LES

4 HFSTI*

5

50,000

LFSTI* 0.45

6 HFSTI*

7

HFSTI 0.56

8 0.75

4

NASA/CR—2012-217415 155
z+ ~ 15-40 (Piomelli and Chasnov [19]). Table 4 shows more 

specifics of Grid#4. Time step size of 0.001 s was used for 2D 

URANS simulations and 0.0001 s for LES. 

Statistical averages of the pressure coefficient plotted 

versus dimensionless distance along both the suction and 

pressure side of an airfoil are presented on Figure 2 for 

Re=50,000, LFSTI. On the figure, the peak of Cp on the 

suction side of the airfoil is located between s/Ls 0.4 and 0.5 

in both CFD and experiment. A plateau in Cp downstream of

the Cp peak indicates flow separation.

 Table 3 Grids used in this investigation

Grid # Size 

(Cells)

Number of 

grids in z 

direction

y+ z+ x+

1 

(2-D)

60,600 NA 0.132 NA < 28

2*

(2-D)

122,000 NA 0.115 NA < 10

3

(2-D)

200,000 NA 0.113 NA < 4

4*

(3-D)

1.7 

million

32 0.117 <1 < 18

Table 4: Specification of computational Grid#4

In cases without wakes (Ibrahim et al. [14]), this plateau 

is large and Cp remains almost constant downstream of the 

peak in both CFD and experiment. This corresponds to a large 

separation without reattachment. In cases with an upstream 

rod the plateau in Cp on the suction side is much shorter 

compared with the cases without a rod and the value of Cp

decreases downstream of the plateau. This indicates a smaller 

separation with reattachment at the end of the plateau in Cp. 

For cases with an upstream rod the comparison between the 

2D and the 3D Cp profiles shows a similar trend of separation 

and reattachment. However, in 2D simulation values of the Cp

peak are under-predicted, compared with experiment. Results 

of the 3D LES simulation show excellent agreement with 

experiment in values and the slope of the Cp profile at all 

airfoil locations. Computational velocity profiles at the inlet 

into the cascade are compared with experimental ones in Fig. 

3. The wake from a rod predicted in 3D LES simulation is in a

reasonable agreement with experimental data for the same 

freestream turbulence conditions. Velocities in the wake are 

over-predicted in the 2D simulation between /L -0.1 and 0.1.

Wake Characteristics

Several studies have compared rod wakes to airfoil 

wakes. Pluim et al. [2] tested rods of various shapes and 

provided references to earlier studies. They found that the 

wakes from wedge shaped rods agreed best with L1A airfoil 

wakes, but cylindrical rods also provided a reasonable 

approximation. 

In the experiment, the moving rod wakes were measured 

by placing a stationary hot-wire midway between blades B4 

and B5 (see Fig. 1) in the plane of the leading edges. The 

moving rods traversed 0.54Cx upstream of the hot-wire. The 

data from the probe were phase-averaged on the rod passing 

cycle to determine the mean and fluctuating velocity as 

functions of time. For comparison, the time averaged velocity 

profile was taken 0.54Cx downstream of the rod, as shown in 

Fig. 3. All curves were normalized using the average local 

mean velocity. The agreement between CFD and the 

experiment is reasonable. 

Figure 4 shows instantaneous vorticity contours for the 

flow behind the rod in the simulations, for Re 50,000. The 

Reynolds number based on the rod diameter for these cases is 

about 594 and the Strouhal number is 0.117 (based on the 

cylinder shedding frequency) which matched the empirical 

correlation obtained by Norberg [20]. The 2D simulation (not 

shown here) showed large organized coherent structures 

(Karman vortex shedding) consistent with a laminar-like flow. 

The 3D, LES simulation (Fig. 4a), on the other hand, shows 

the wake structures with small-scale eddies as expected in 

reality. Fig. 4b shows the initial attempt to run a URANS, 3D 

model at the same time step size as 2D (t = 0.001 s). The 

figure shows that this time step was not sufficiently fine 

enough to capture the vortex shedding Lardeau [12]. Thus the 

same size time step used for the LES calculations was 

necessary to get satisfactory results, losing the computational 

advantage of using a URANS model with a larger time step. 

Thus for the rest of the paper, all 3D cases were run using 

LES.

Number of cells (million) 1.7

Number of nodes on the suction surface 292

Number of nodes on the pressure surface 240

Number of nodes in span direction 32

y+ 0.117

z+ < 1

x+ < 18

Distance from inlet boundary to the leading edge 3.8 Cx

Distance from the trailing edge to the outlet 

boundary 

1.9 Cx

5

NASA/CR—2012-217415 156
Re=25,000, F=0.45

Figures 5 and 6 show the time averaged Cp profiles and 

total pressure loss coefficient () profiles for the 2D and 3D 

cases, compared against the experimental data. Here it can be 

seen that in the 2D case, the lack of ability to accurately 

predict the magnitude of the Cp peak, along with an overprediction in the separation bubble size (shown by a large 

plateau in the Cp values), result in greater prediction in the 

peak of the Total Pressure Loss Coefficient () and a sharper 

drop in the profile. 

For the 3D cases, both the LFSTI and HFSTI cases match 

reasonably good to the experimental data, however the lack of 

ability to locate the separation and reattachment points lead to 

a larger predicted separation bubble in the LFSTI case and a 

smaller predicted one in the HFSTI case. This affects , 

showing a peak loss greater than the experimental data for the 

LFSTI case and a peak loss less than the experimental data for 

the HFSTI case. There is also a much sharper drop in the 

profile due to a sharp drop in the Cp profile after reattachment.

Reynolds 25,000 cases are difficult to document due to 

the small change in pressure associated with the Cp and total 

pressure loss coefficients, thus leading to significant 

uncertainty in both the experimental and computational cases. 

Higher accuracy is expected as the Reynolds number is 

increased. 

Re=50,000, F=0.45

Figures 7 and 8 show the results of the time averaged Cp

profiles and Total Pressure Loss Coefficient () profiles for 

the 2D URANS and 3D LES cases compared against the 

experimental data. Once again, in the 2D URANS simulation 

the magnitude of the Cp peak is under predicted, however the 

size of the bubble is much more accurate compared to the 

Re=25,000 case, and is consistent with both the 3D LES case 

and the experimental data. This results in a ψ peak value that 

is consistent with the experimental data, however values of ψ 

between -0.2 and 0.5 /L are over-predicted compared with 

experiment. The poorer agreement between the experiment 

and predicted ψ profiles (compared to Cp profiles) could be 

attributed to: 1) the mesh structure used in this wake region 

and 2) the turbulence model applied.

Results of the 3D LES LFSTI and HFSTI cases match 

the experimental data reasonably well in both the magnitude 

of the Cp profile as well as in predicting the locations of 

separation and reattachment. In these cases the magnitude of 

the ψ peak as well as an overall shape of the ψ profile for /L

between -0.5 and 0.2 agree with experimental data. 

To provide further insight into the wake influence on the 

suction surface boundary layer separation and reattachment 

data from the 3D LES simulations under LFSTI and HFSTI 

conditions are analyzed below. 

Velocity measurements were conducted at 5 different 

stations downstream of the location of the Cp peak (see Table 

5). 

Table 5: Velocity profile measurement stations

Station 1 2 3 4 5

s/Ls 0.59 0.69 0.78 0.88 0.97

x/Cx 0.72 0.80 0.86 0.92 0.97

Figure 9 shows the velocity profiles at the 5 streamwise 

locations described in Table 5 for the CFD Re=50,000, HFSTI 

case with a rod to blade spacing of 1. Experimental velocity 

profiles from the case with the rod to blade spacing of 1.6 are 

shown for comparison. This comparison is justified, since 

pressure coefficients obtained from experiment with rod to 

blade spacings of 1.6 and 1 are similar and also similar to the 

Cp profile predicted in CFD (spacing 1). Predicted mean 

velocity profiles, shown on Fig. 10a, agree with experimental 

ones well. A small separation bubble is predicted in CFD at 

stations 1 and 2 which is indicated by the negative velocities 

near the wall. Reattachment is predicted at stations 3-5. This is 

consistent with separation and reattachment locations derived 

from the Cp profile shown on Fig. 7. It is more difficult to 

determine separation and reattachment locations from 

experimental velocity profiles, since the hot wire anemometry 

method, used in the experiment to measure velocities, does not 

allow identifying flow direction. Figure 9b shows the RMS 

velocity profiles for the same cases. Here, the CFD shows the 

same general profile, matching well with the data; however it 

is consistently greater in u̒/Ue, even in the freestream. This 

may be the result of the fact that the wake profile seen in Fig.

3 shows a wake with a greater deficit and larger width than the 

experiment, thus resulting in the increase in turbulence levels 

in the passage. This large width in the wake indicates that 

Direct Numerical Simulation (DNS) might be needed to 

resolve turbulence in the wake of the rod and in the cascade, 

however due to large computational times associated with 

DNS this would be prohibitive.

Figures 10 and 11 show the same mean Cp profiles for the 

LFSTI and HFSTI cases shown earlier in Fig. 8, however this 

time they are plotted along with the mean coefficient of 

friction (Cf) profiles. Marked on these profiles are dotted lines 

showing the indications of separation, transition, and 

reattachment. It can be seen that in the Cf profile, where the 

plateau in the Cp profile begins, thus indicating separation, an 

inflection in the Cf curve is also seen. This inflection reduces 

the slope of the curve slightly, allowing for Cfto more slowly 

approach a point where it hits zero and then equally slowly 

starts to increase again. At a point just before reattachment,

indicated by the Cp profile, there is another inflection in the Cf

curve resulting in a more increased slope and a more rapid 

6

NASA/CR—2012-217415 157
increase in Cf

. This is the indication of transition beginning to 

occur. Finally, at the point of reattachment, the Cf curve 

indicates another inflection right before it reaches its peak. 

In Figures 12-1 through 12-4, instantaneous results for the 

Re=50,000, HFSTI case are shown for one full wake passing 

cycle. The Cf and Cp profiles are visualized by plotting their 

values on separate y-axis while the Q-Criterion contours 

colored by Cp values are shown beneath to give a visualization 

of the location of the wake in the passage and its effect on the 

pressure profile. 

Figure 12-1 shows that as the wake has collapsed over the 

Cp peak and suppression of the separation bubble begins, the 

bubble is rolled downstream with the Cp peak, with the 

associated peak Cf being compressed and increasing in 

magnitude. Trailing behind this spike in turbulence is a 

calmed region which has near zero Cf, but is resistant to 

separation. These peaks as well as the trailing calmed region 

were predicted in Hodson and Howell [1]. The contour plot 

shows the elongated wake structures forcing the separation 

bubble downstream. 

Figure 12-2 shows 30% through a cycle, the separation 

bubble has been rolled nearly all the way out of the cascade 

passage and has been removed from the surface. The curves 

show a large amount of turbulence still present on the aft 

portion of the suction surface; however the once calmed 

region is starting to show localized points of separation. The 

Cp peak is nearly fully reestablished as another wake 

approaches upstream. 

Figure 12-3 shows 60% of a cycle at which point the Cp

peak has a very dominant profile. The separation bubble is 

nearly fully re-grown, with the presence of the wake in the 

forward part of the passage showing no effect on the boundary 

layer of the suction surface.

Figure 12-4 shows that at 90% through the cycle, the 

wake is beginning to collapse over the Cp peak and affect the 

separation bubble. The bubble is being rolled up, with the 

new peak in Cp and a peak in Cf developed. 

Re=50,000, F=0.56, F=0.75

For a further analysis of the effect of the wake, the 

dimensionless wake passing frequency, F, was increased by 

increasing the speed of the rod. Figure 13 shows that the 

initial 25% increase in F (to the value of F=0.56) resulted in a 

significant reduction in the size of the separation bubble as 

compared to the original F=0.45 case, nearly suppressing 

separation completely. The result in the Total Pressure Loss 

Coefficient (ψ) seen in Fig. 14 shows a peak loss also greatly 

reduced. However, when trying to further increase the rod 

speed an additional 33% (to the value of F=0.75), the resulting 

profiles are nearly identical to the F=0.56 case. The size of 

the separation bubble and thus the resulting ψ profile are 

nearly identical. This illustrates that a wake can suppress 

separation given a high enough passing frequency, however 

any further increase in the wake frequency will have little 

effect on the flow separation on the suction surface. 

Summary and Conclusions 

In the present study 2D URANS and 3D LES simulations 

were performed in order to investigate the effect of the wake 

from an upstream moving rod on the flow separation and 

reattachment on the highly loaded airfoil L1A. Computational 

results were compared with experimental data [4]. Results of 

2D URANS simulations matched the experimental trend of 

interaction between the wake and the boundary layer on the 

suction surface of the airfoil with the mean pressure 

coefficient profiles (Cp) indicating separation followed by 

reattachment. These trends were only marginally accurate in 

comparison to the experimental data for both Re=25,000 and 

Re=50,000 cases, since the predicted magnitude of the Cp

profile and the location of reattachment were not very 

consistent with the experimental data. This was mainly 

attributed to the lack of ability to take into account the three 

dimensional structure of a wake and its subsequent interaction 

with the boundary layer. Thus the CFD mesh was then 

projected into a 3D domain and the same URANS (TransSST) model was run at the same time step as in the 2D cases. 

These 3D URANS cases were unable to capture any wake 

vortex shedding of the rod, which had been seen in cases run 

by Ladreau [12]. It was determined that in order to capture 

the wake, the time step size would have to be reduced to levels 

associated with LES turbulence, thus the remainder of the 

cases were run using LES since the computational cost savings 

of using a URANS model with a larger time step were 

diminished.

For Re=50,000 3D LES predictions were in a reasonably 

good agreement with experimental data in both pressure and 

velocity profiles on the airfoil. For Re=25,000 the predictions 

were not as good. Instantaneous profiles of Cp and Cf were 

analyzed, in order to explain the effect of the wake, as it 

passes through the cascade passage, on the suction surface

boundary layer. These instantaneous results showed that as a 

wake reaches the location of the Cp peak, it collapses down, 

compressing the peak Cp value and eventually rolling it up 

into three defined peaks. These peaks first experience a shear, 

stretching effect and eventually begin to roll down the aft side 

of the suction surface, remaining attached to the surface and 

being further compressed and increasing in Cp magnitude. 

The corresponding effect on the Cf values over these events 

show that as the suction bubble is broken down and 

compressed, the three defined Cp peaks are associated with 

defined peaks in the Cf values, thus showing a high amount of 

turbulence in this region. Trailing behind this turbulent peak 

is the predicted calmed region described by Hodson and 

Howell [1]. This calmed region remains near zero in Cf value, 

7

NASA/CR—2012-217415 158
but shows resistance to separation as long as the separation 

bubble remains attached to the surface. As this bubble is 

eventually rolled off the surface, the calmed regions show an 

immediate location of separation with a zero Cf value. This 

separation grows and eventually spans across the width of the 

suction surface as the Cp values start to show a defined 

separation bubble, indicated by a “plateau” in the Cp profile. 

Also shown by the Cf profile is where separation is indicated 

to start by an inflection in the decreasing Cf values. The 

leveling out of this curve as it approaches, hits, and recovers 

from a zero Cf value indicates where the separation bubble 

exists, as seen by the plateau in the Cp profile. As it starts to 

increase out of this zero value, another inflection occurs in the 

curve slightly before reattachment as indicated on the Cp

values. This inflection shows a start of a sharp increase in Cf

and indicates the start of transition in the boundary layer. 

Reattachment occurs just before another peak in the Cf values, 

which then shows turbulence in the flow as it stays attached to 

the trailing edge.

For Re=50,000 the effect of the wake passing frequency 

on the flow separation and reattachment on the airfoils suction 

surface was investigated. Initial increase in dimensionless 

wake frequency, from F = 0.45 to F= 0.56, helped to reduce 

the size of the separation bubble to near nonexistent; however 

a slight indication of one still was present. Additional increase 

in dimensionless wake frequency, from F = 0.56 to F = 0.75,

showed little effect on the airfoil flow. 

Acknowledgements 

This work was sponsored by the National Aeronautics and 

Space Administration. The grant monitor is Dr. James 

Heidmann of the NASA Glenn Research Center. The support 

of the United States Naval Academy Technical Support 

Department Shop and Fluids Laboratory is greatly 

appreciated. We greatly appreciate the computer time 

provided for us by the Ohio Super Computer (OSC). The OSC 

Computer Cluster has been made available as part of the 

Center’s mission to support Ohio Universities.

References 

[1] Hodson, H.P., and Howell, R.J., 2005, “Blade Row

Interactions, Transition, and High-Lift Airfoils in LowPressure Turbine,” Annual Review of Fluid Mechanics, 

37, 71-98.

[2] Pluim, J., Memory, C., Bons, J., and Chen, J.P., 

2009, “Designing a High Fidelity Wake Simulator for 

Research Using Linear Cascades,” ASME Paper 

GT2009-52276. 

[3] Bons, J.P., Pluim, J., Gompertz, K. and Bloxham, 

M., 2008, “The Application of Flow Control to an AftLoaded Low Pressure Turbine Cascade with Unsteady 

Wakes”, ASME Paper GT2008-50864. 

[4] Volino, R.J., 2012, “Effect of Unsteady Wakes on 

Boundary Layer Separation on a Very High Lift Low 

Pressure Turbine Airfoil,” ASME J. Turbomachinery, 

134, paper 011011. 

[5] Sarkar, S., 2009, “Influent of Wake Structure on 

Unsteady Flow in a Low Pressure Turbine Blade 

Passage," ASME J. Turbomachinery, 131, paper 041016. 

[6] Stieger, R., Hollis, D. and Hodson, H., 2003, 

“Unsteady Surface Pressures Due to Wake Induced 

Transition in Laminar Separation bubble on a LP 

Turbine Cascade”, ASME Paper GT2003-38303. 

[7] Suzen, Y.B. and Huang, P.G., 2005, “Numerical 

Simulation of Unsteady Wake/Blade Interactions in 

Low-Pressure Turbine Flows Using and Intermittency 

Transport Equation”, ASME J. Turbomachinery, 127, 

431-444. 

[8] Kaszeta, R. W., Simon, T. W., and Ashpis, D. E., 

2001, “Experimental Investigation of Transition to 

Turbulence as Affected by Passing Wakes,” ASME 

Paper 2001-GT-0195. 

[9] Kaszeta, R. W., and Simon, T. W., 2002, 

“Experimental Investigation of Transition to Turbulence 

as Affected by Passing Wakes,” NASA Contractor 

Report, NASA-CR-2002-212104, Cleveland. 

[10] Stieger, R. D., 2002, “The Effects of Wakes on 

Separating Boundary Layers in Low-Pressure Turbines,” 

Ph.D. Dissertation, Cambridge University, Eng. Dept., 

Cambridge, UK 

[11] Rodi, W., 2006, “DNS and LES of some 

Engineering Flows”, Fluid Dynamics Research, 38, 145-

173.

[12] Lardeau, S., 2005, “Unsteady RANS modeling of 

wake-blade interaction: computational requirements and 

limitations,” Computers and Fluids, 34, 3-21.

[13] ANSYS FLUENT Documentation, 2009. 

[14] Ibrahim, M.B., Kartuzova, O., Doucet, D.J. and 

Volino, R.J., 2010, “LES Flow Control Simulations For 

Highly Loaded Low Pressure Turbine Airfoil (L1A) 

Using Pulsed Vortex Generator Jets”, ASME Paper 

GT2010-23015. 

[15] Menter, F.R., Langtry, R.B., Likki, S.R., Suzen, 

Y.B., Huang, P.G., and Völker, S., 2006, “A Correlation 

based Transition Model using Local Variables Part 1-

Model Formulation”, ASME J. Turbomachinery, 128, 

413-422. 

[16] Ibrahim, M.B., Kartuzova, O. and Volino, R.J., 

2008, “Experimental and Computational Investigations 

of Separation and Transition on a Highly Loaded Low 

Pressure Turbine Airfoil: Part 1 – Low Freestream 

Turbulence Intensity”, ASME Paper IMECE2008-68879. 

8

NASA/CR—2012-217415 159
[17] Kim, W.W. and Menon, S., 1997, “Application of 

the localized dynamic subgrid-scale model to turbulent 

wall-bounded flows,” AIAA Paper AIAA-97-0210. 

[18] Kim S.E., 2004, “Large eddy simulation using 

unstructured meshes and dynamic subgridscale 

turbulence models,” AIAA Paper AIAA-2004-2548. 

[19] Piomelli, U. and Chasnov, J.R., 1995, “Large-Eddy 

Simulations: Theory and Applications”, Turbulence and

Transition Modeling, lecture notes from the

ERCOFTAC/IUTAM summer school held in Stockholm, 

12-20 June, 1995.

[20] Norberg, C., 2003, “Fluctuating Lift on a Circular 

Cylinder: Review and New Measurements”, J. Fluid 

Struct., 17, 57-96.

Figure 2. Cp profiles for Re=50,000, LFSTI: Comparison 

between Grid and Turbulence Model against Experimental 

data.

Figure 3. Rod Wake Velocity Profile at Inlet Cascade 

Plane, Re=50,000: Comparison between 3D (LES) LFSTI, 

3D (LES) HFSTI, and Experimental Data.

a) b)

Vorticity (m/s)

 5,500 3,850 1,920 825 0

Figure 4. Vorticity Contours for Re = 50,0000: Comparison between a) 3D (LES) b) 3D (Trans-SST).

0.00

0.50

1.00

1.50

2.00

2.50

0.00 0.20 0.40 0.60 0.80 1.00

Cp

s/Ls

Exp. No Wake Exp. 1LØ

CFD No Wake CFD - 3D - LES

CFD - 2D - Trans-SST

0.7

0.8

0.9

1

1.1

1.2

-0.5 -0.3 -0.1 0.1 0.3 0.5

U/Ue

φ/L

CFD 3D 1Lφ LFSTI CFD 3D 1Lφ HFSTI

CFD 2D 1Lφ LFSTI Exp. 1.6Lφ

9

NASA/CR—2012-217415 160
Figure 5. Cp profiles for Re=25,000, F=0.45: Comparison 

between LFSTI, HFSTI, 3D (LES), 2D (URANS) and 

Experimental data.

Figure 6. Total Pressure Loss Coefficient for Re=25,000, 

F=0.45, 0.68Cx Downstream: Comparison between LFSTI, 

HFSTI, 3D (LES), 2D (URANS) and Experimental data.

Figure 7. Cp profiles for Re=50,000, F=0.45: Comparison 

between LFSTI, HFSTI, 3D (LES), 2D (URANS) and 

Experimental data.

Figure 8. Total Pressure Loss Coefficient for Re=50,000, 

F=0.45, 0.68Cx Downstream: Comparison between LFSTI, 

HFSTI, 3D (LES), 2D (URANS) and Experimental data.

0.00

0.50

1.00

1.50

2.00

2.50

0.00 0.20 0.40 0.60 0.80 1.00

Cp

s/Ls

CFD 3D HFSTI Series10

CFD 3D LFSTI Series4

CFD 2D LFSTI Exp. HFSTI 1.6Lф

-0.2

0.0

0.2

0.4

0.6

0.8

1.0

1.2

-0.5 -0.4 -0.3 -0.2 -0.1 0.0 0.1 0.2 0.3 0.4 0.5

ψ

φ/Lφ

CFD 3D HFSTI Exp. HFSTI

CFD 3D LFSTI Exp LFSTI

CFD 2D LFSTI

0.00

0.50

1.00

1.50

2.00

2.50

0.00 0.20 0.40 0.60 0.80 1.00

Cp

s/Ls

CFD 3D HFSTI Exp HFSTI

CFD 3D LFSTI Series8

CFD 2D LFSTI Exp. HFSTI 1.6Lф

0.0

0.2

0.4

0.6

0.8

1.0

-0.5 -0.4 -0.3 -0.2 -0.1 0.0 0.1 0.2 0.3 0.4 0.5

ψ

φ/Lφ

CFD 3D HFSTI Exp. HFSTI

CFD 3D LFSTI Exp LFSTI

CFD 2D LFSTI

Exp. LFSTI

Exp. HFSTI Exp. LFSTI

10 NASA/CR—2012-217415 161
a)

b)

Figure 9. Time averaged velocity profiles at five streamwise stations for Re=50,000, HFSTI: Comparison between CFD (LES) 

and Experimental data: a - Mean, b – RMS.

0

0.05

0.1

0.15

0.2

0.25

-0.5 0 0.5 1 1.5

U/Ue

s/Ls = 0.59

CFD

Exp. 1.6

-0.5 0 0.5 1 1.5

U/Ue

s/Ls = 0.69

-0.5 0 0.5 1 1.5

U/Ue

s/Ls = 0.78

-0.5 0 0.5 1 1.5

U/Ue

s/Ls = 0.88

-0.5 0 0.5 1 1.5

U/Ue

s/Ls = 0.97

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0 0.1 0.2 0.3 0.4

u'/Ue

s/Ls =0.59

0 0.1 0.2 0.3 0.4

u'/Ue

s/Ls = 0.69

0 0.1 0.2 0.3 0.4

u'/Ue

s/Ls = 0.78

0 0.1 0.2 0.3 0.4

u'/Ue

s/Ls = 0.88

0 0.1 0.2 0.3 0.4

u'/Ue

s/Ls = 0.97

Figure 10. Comparison of Mean Cp and Cf on suction 

surface for Re=50,000, LFSTI, F=0.45.

Figure 11. Comparison of Mean Cp and Cf on suction 

surface for Re=50,000, HFSTI, F=0.45.

0.00 0.20 0.40 0.60 0.80 1.00

0.000

0.005

0.010

0.015

0.020

0.025

0.030

0.00

0.50

1.00

1.50

2.00

2.50

Cf

Cp

s/Ls

Cp Cf

0.00 0.20 0.40 0.60 0.80 1.00

0.000

0.005

0.010

0.015

0.020

0.025

0.030

0.00

0.50

1.00

1.50

2.00

2.50

C

C

f

p

s/Ls

Cp Cf

11 NASA/CR—2012-217415 162
 

Cp

Figure 12-1. Comparison of Re=50,000, LFSTI, F=0.45: Instantaneous 

Cp and Cf on suction surface and Q-criterion contours colored by Cp

values at time t/T = 0.

Cp

Figure 12-2. Comparison of Re=50,000, LFSTI

Cp and Cf on suction surface and Q-criterion 

values at time t/T = 0.30.

0.00 0.20 0.40 0.60 0.80 1.00

0.050

0.040

0.030

0.020

0.010

0.000

2.50

2.00

1.50

1.00

0.50

0.00

f C

p C

s/Ls

Cp Cf

0.00 0.20 0.40 0.60 0.80 

2.50

2.00

1.50

1.00

0.50

0.00

p C

s/Ls

Cp Cf

12

NASA/CR—2012-217415 163
Cp

Figure 12-3. Comparison of Re=50,000, LFSTI, F=0.45: Instantaneous 

Cp and Cf on suction surface and Q-criterion contours colored by Cp

values at time t/T = 0.60.

Cp

Figure 12-4. Comparison of Re=50,000, LFSTI

Cp and Cf on suction surface and Q-criterion 

values at time t/T = 0.90.

0.00 0.20 0.40 0.60 0.80 1.00

0.050

0.040

0.030

0.020

0.010

0.000

2.50

2.00

1.50

1.00

0.50

0.00

f C

p C

s/Ls

Cp Cf

0.00 0.20 0.40 0.60 0.80 

2.50

2.00

1.50

1.00

0.50

0.00

p C

s/Ls

Cp Cf

13

NASA/CR—2012-217415 164
Figure 13. Mean Cp on suction surface for Re=50,000, HFSTI: 

Comparison between CFD F=0.45, F=0.56, and F=0.75.

Figure 14. Total Pressure Loss Coefficient for R

Comparison between CFD F=0.45, F=0.56, and 

0.00 0.20 0.40 0.60 0.80 1.00

2.50

2.00

1.50

1.00

0.50

0.00

p C

s/Ls

Cp F=0.917 Cp F=0.548 Cp F=0.689

1.0

0.8

0.6

0.4

0.2

0.0

-0.5 -0.4 -0.3 -0.2 -0.1 0.0 0.1 

ψ

φ/Lφ

CFD F=0.45 CFD F=0.56 

14

NASA/CR—2012-217415 165
Ralph J. Volino

Department of Mechanical Engineering,

United States Naval Academy,

Annapolis, MD 21402

e-mail: volino@usna.edu

Separated Flow Measurements

on a Highly Loaded Low-Pressure

Turbine Airfoil

Boundary layer separation, transition, and reattachment have been studied on a new,

very high lift, low-pressure turbine airfoil. Experiments were done under low freestream

turbulence conditions on a linear cascade in a low speed wind tunnel. Pressure surveys

on the airfoil surface and downstream total pressure loss surveys were documented.

Velocity profiles were acquired in the suction side boundary layer at several streamwise

locations using hot-wire anemometry. Cases were considered at Reynolds numbers

(based on the suction surface length and the nominal exit velocity from the cascade)

ranging from 25,000 to 330,000. In all cases, the boundary layer separated, but at high

Reynolds number the separation bubble remained very thin and quickly reattached after

transition to turbulence. In the low Reynolds number cases, the boundary layer separated

and did not reattach, even when transition occurred. This behavior contrasts with previous research on other airfoils, in which transition, if it occurred, always induced reattachment, regardless of Reynolds number. DOI: 10.1115/1.3104608

1 Introduction

Boundary layer separation can lead to partial loss of lift and

higher aerodynamic losses on low-pressure turbine LPT airfoils

e.g., Refs. 1–3. As designers impose higher loading to improve

efficiency and lower cost, the associated strong adverse pressure

gradients on the suction side of the airfoil can exacerbate separation problems. The problem is particularly relevant in aircraft engines due to the lower density and therefore lower Reynolds numbers at altitude. A component efficiency drop of 2% may occur

between takeoff and cruise in large commercial transport engines,

and the difference could be as large as 7% in smaller engines

operating at higher altitudes 4,5. Prediction and control of suction side separation, without sacrifice of the benefits of higher

loading, are therefore crucial for improved engine design.

Separation and separated flow transition, which can lead to

boundary layer reattachment, have received considerable attention. Studies have included flows over flat plates subject to pressure gradients similar to those on the suction side of LPT airfoils,

and flows over airfoils either in single passage facilities or multiblade cascades. Some have considered steady inflow conditions,

while others have included the effect of unsteady wakes. Recent

examples include Refs. 6–12. Volino 13 provided a review of

some earlier studies. In general, previous work shows that the

strong acceleration on the leading section of the airfoil keeps the

boundary layer thin and laminar, even in the presence of elevated

freestream turbulence. When separation does occur, it is usually

just downstream of the suction peak. If transition then occurs in

the shear layer over the separation bubble, it is typically rapid and

causes the boundary layer to reattach 13,14. Transition is dependent on Reynolds number, freestream turbulence level, and the

surface roughness conditions upstream of the separation point.

Several correlations for separated flow transition have been developed based on experimental data, including those of Mayle 2,

Hatman and Wang 15, Davis et al. 16, Yaras 17, Volino and

Bohl 18, and Praisner and Clark 19.

The advances in separation understanding and prediction have

led to attempts at separation control. Zhang et al. 20, Bohl and

Volino 21, Volino 22, and others provided examples using passive devices such as boundary layer trips. Others have used active

devices such as vortex generator jets e.g., Refs. 4,23 or plasma

devices e.g., Ref. 24.

Another way to improve performance is to design airfoils with

pressure gradients more resistant to separation, as described by

Praisner and Clark 19. Forward loading, for example, makes

airfoils more separation resistant by extending the adverse pressure gradient on the aft portion of the suction side over a longer

distance. This reduces the local pressure gradient at all locations,

making separation less likely. If separation does occur, forward

loading provides a longer distance along the airfoil surface for

reattachment. Forward loading has some disadvantages, however.

As noted by Zhang et al. 20, the longer region of turbulent flow

on a forward loaded airfoil can lead to increased profile losses.

Forward loading also creates longer regions of strong pressure

gradient on the endwalls, which can produce stronger secondary

flows and losses. If flow control were incorporated in the design

of an advanced airfoil, as discussed by Bons et al. 25, it might be

possible to produce an aft loaded airfoil that was resistant to separation and had low profile and secondary loss characteristics over

a range of Reynolds numbers.

The objective of the present study is to document the flow over

a very highly loaded LPT airfoil. The airfoil chosen was designed

at the Air Force Research Laboratory AFRL, and is designated

the L1A. It is available on a limited basis to U.S. researchers from

Ref. 26. Dimensions of the L1A as used in the present experiments are given in Table 1. Based on the design calculations of

Clark 26, the L1A has 10% higher loading than the “ultrahigh

lift” airfoils described by Zhang and Hodson 27, and 17% higher

loading than the Pack B airfoil considered in several studies such

as Refs. 13,14,4. The design calculations indicate that the Zweifel coefficient increases from about 1.15 on the Pack B to about

1.35 on the L1A. If the definition for the Zweifel coefficient is

taken as

Zw = 2 cos2 2L/Cxtan 1 + tan 2 1

as given by Lakshminarayana 28, Zw= 1.08 for the Pack B in

agreement with the value given by McAuliffe and Sjolander 29

and Zw= 1.23 for the L1A. The L1A has the same inlet and exit

flow angles and loading as the L1M airfoil used by Bons et al.

25. The L1M is a midchord loaded design and is resistant to

Contributed by the International Gas Turbine Institute of ASME for publication in

the JOURNAL OF TURBOMACHINERY. Manuscript received July 7, 2008, final manuscript

received: February 2, 2009; published online September 16, 2009. Review conducted

by David Wisler. Paper presented at the ASME Turbo Expo 2008: Land, Sea and Air

GT2008, Berlin, Germany, June 9–13, 2008.

Journal of Turbomachinery Copyright © 2010 by ASME JANUARY 2010, Vol. 132 / 011007-1

Downloaded 11 Jan 2011 to 131.122.82.80. Redistribution subject to ASME license or copyright; see http://www.asme.org/terms/Terms_Use.cfm

NASA/CR—2012-217415 166
separation even at very low Reynolds numbers. While the L1M

and L1A were designed with the same methodology, the L1A is

aft loaded, and is therefore expected to be more prone to separation. It is expected to be a good test airfoil for future flow control

work, in which separation may be suppressed while maintaining

the benefits of both very high loading and aft loading. Figure 1

shows the inviscid flow pressure profiles for the L1A, L1M, and

Pack B airfoils. The inviscid code used for the computations was

adapted from Ref. 30. Figure 1a shows the pressure coefficient

Cp as a function of dimensionless axial position, x/Cx. These

coordinates are useful for comparing the lift generated by the

airfoils. Figure 1b shows Cp as a function of the streamwise

distance along the airfoil surface, s/Ls, where Ls is the wetted

surface length along the suction side. These coordinates are most

useful for explaining the boundary layer development and are

used to present the results below. The higher lift of the L1A and

L1M compared with the Pack B is clear in Fig. 1. Also clear is the

stronger adverse pressure gradient of the L1A on the downstream

region of the suction side. The pressure gradient is shown again in

Fig. 2 as the product of the local acceleration parameter K and the

Reynolds number. The Reynolds number, Re, is based on Ls and

the nominal exit velocity from the cascade Ue computed using

the inlet velocity and the design inlet and exit flow angles. The

parameter K is inversely proportional to Re, so K Re is independent of Reynolds number. The negative K values downstream of

s/Ls= 0.7 are about twice as strong for the L1A than for the other

two airfoils. This suggests more of a tendency for boundary layer

separation and possibly less of a tendency for the flow to reattach.

2 Experimental Facility and Measurements

Experiments were conducted in a closed loop wind tunnel with

a linear cascade in one corner of the loop. An axial fan sends air

through turning vanes in the first corner of the tunnel and then

through a heat exchanger. Following the heat exchanger is the

second turn and a 6.25:1 area ratio contraction. Following the

contraction is a 0.46 m square by 1.1 m long test section, which

was not used in the present study. Following this test section is a

2.4 m long diffuser, which expands the flow area to 0.75

0.75 m2. A fine screen with 0.12 mm mesh thickness, 0.42 mm

mesh spacing and 49% blockage is located in the exit plane of the

diffuser. The screen breaks up the boundary layers, which form on

the diffuser walls and provides a uniform flow into the cascade

test section. A seven blade cascade is located in the wind tunnel’s

third turn, as shown in Fig. 3. A generic airfoil shape is shown in

the figure. The flow continues from the cascade into the tunnel’s

fourth turn, then through a second diffuser before returning to the

fan.

The freestream turbulence entering the cascade was measured

with a cross-wire probe positioned just upstream of the center

blade. The streamwise velocity component had a turbulence intensity, TI, of 0.8%, while the component intensity in the direction

across the cascade was 0.5%. The spanwise component intensity

was assumed to also be 0.5% based on the symmetry of the wind

tunnel. The combined freestream turbulence intensity based on all

Table 1 Cascade parameters

Axial chord, Cx

mm

True chord

mm

Pitch, L

mm

Span

mm

Suction side, Ls

mm

Inlet flow angle

deg

Exit flow angle

deg

134 146 136 724 203 35 60

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

x/Cx

Cp

Pak B

L1M

L1A

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

Pak B

L1M

L1A

(b)

(a)

Fig. 1 Comparison of inviscid pressure profiles for the Pack

B, L1M, and L1A airfoils: „a... Cp versus axial position; „b... Cp

versus streamwise location

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 −2

−1

0

1

2

3

4

5

s/Ls

KRe

Pak B

L1M

L1A

Fig. 2 Acceleration, K Re, versus streamwise location for

Pack B, L1M, and L1A airfoils based on inviscid solution

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NASA/CR—2012-217415 167
three components was 0.6%. Spectral measurements showed that

80% of the energy in the streamwise fluctuations was due to the

unsteadiness at dimensionless frequencies fLs/Ue2, where f is

the frequency in hertz. Approximately 70% of the energy in the

cross stream components was also below this frequency. If the

turbulence is high pass filtered to remove this unsteadiness, the

streamwise and cross stream component intensities are 0.36% and

0.27%, respectively, for an overall freestream turbulence intensity

of 0.3%. The integral length scale of the freestream turbulence is

6.3 cm in the streamwise direction and 6.7 cm in the other directions. While such low freestream turbulence and large length

scales are not representative of engine conditions, they are still of

interest. In zero or favorable pressure gradient boundary layers,

high TI can cause bypass transition, but under adverse pressure

gradients, natural transition appears to play a role at all TI levels.

Hughes and Walker 31, for example, observed evidence of

Tollmien–Schlichting TS waves in cases with TI as high as 8%.

Volino 14 also saw evidence of TS waves in both high and low

TI cases. Low TI cases provide a somewhat simpler environment

for explaining the already complex phenomenon of separated flow

transition. The present low TI cases will also serve as comparison

cases for future work with elevated TI. Higher TI and length

scales closer to the boundary layer thickness are expected to hasten the transition process.

The blades in the cascade were machined from high density

foam, which has a consistency much like hard wood. The center

blade, designated B4 in Fig. 3, contains pressure taps near the

spanwise centerline. A tailboard, shown in Fig. 3, was needed to

produce the correct exit flow angle from the cascade. Its position

was set to produce periodicity at high Reynolds numbers. A tailboard on the opposite side of the cascade and inlet guide vanes

were found to be unnecessary. To produce the correct approach

flow to the end blades B1 and B7, the amount of flow escaping

around the two ends of the cascade was controlled with the flaps

shown in Fig. 3. The flap positions were set using a wool tuft

upstream of each blade to check that the incoming flow approached the stagnation points with the correct angle. The inlet

flow angle at the center of the cascade was also checked with a

three-hole pressure probe and found to be within 2 deg of the

design angle. At high Reynolds numbers, the approach velocity to

the middle four passages was measured to be uniform to within

6%, and the difference between any two adjacent passages was

within 3%. At low Reynolds numbers, slightly more variation was

observed, but the approach velocity to the middle two passages

still agreed to within 5%. Good periodicity at high Reynolds numbers was also observed in the exit flow from the cascade, as evidenced by suction side velocity profiles acquired near the trailing

edge of blades B2–B6 and by total pressure loss surveys, which

are shown below. At low Reynolds numbers, when significant

separation bubbles were present, the periodicity was not as good

due to the suppression of the separation bubble thickness on the

blades closest to the tailboard. This is an unavoidable result when

using a finite linear cascade to study separated flow. It is considered acceptable for the present facility, since its intended purpose

is for the study of flow control, which if successful should suppress separation on all blades, thereby restoring periodicity even

at low Reynolds numbers.

2.1 Measurements. Pressure surveys were made using a pressure transducer 0–870 Pa range Validyne transducer. Stagnation

pressure was measured with a pitot tube upstream of the cascade.

Static pressure taps were located in the center blade as noted

above. The uncertainty in the suction side pressure coefficients

was 0.07. Most of this uncertainty was due to the bias error. Stochastic error was minimized by averaging pressure transducer

readings over a 10 s period.

A four component traverse with three linear stages and one

rotating stage was located in the wind tunnel downstream of the

cascade. The traverse produced an acceptably low blockage when

it was located at least two axial chord lengths downstream of the

cascade. Variations in the wind tunnel velocity were less than 2%

as the traverse was moved to various positions. The traverse was

used to hold and move probes for velocity and downstream pressure measurements.

Total pressure losses were documented using a Kiel probe traversed across three blade spacings, 0.63Cx downstream of the

cascade. To compute the mass flow averaged pressure loss, a velocity profile was also acquired along the same line as the total

pressure surveys using a single-sensor hot-film probe.

Velocity profiles on the suction surface were measured at the

six streamwise stations listed in Table 2. All stations are downstream of the inviscid pressure minimum at s/Ls= 0.49. Profiles

were measured near the spanwise centerline of the airfoil. Profiles

were acquired with a hot-wire anemometer AA Lab Systems

model AN-1003 and a single-sensor hot-film probe TSI model

1201-20. The sensor diameter is 51 m, and the active length is

1.02 mm. At each measurement location, data were acquired for

26 s at a 20 kHz sampling rate 219 samples. All raw data were

saved. The high sampling rate provides an essentially continuous

signal, and the long sampling time results in low uncertainty in

both statistical and spectral quantities. Data were acquired at 40

wall normal locations in each profile, extending from the wall to

the freestream, with most points concentrated in the near-wall

region. The probe was positioned as close to tangent to the airfoil

surface as possible at each station, such that the probe body extended downstream of the sensor and the direction of the traverse

was within 5 deg of normal to the surface. In most cases the

closest point to the wall in each profile was within about 0.2 mm

of the wall, which compares to boundary layer thicknesses rangtailboard

flap

flap

screen

B1

B2

B3

B4

B5

B6

B7

Fig. 3 Schematic of linear cascade

Table 2 Velocity profile measurement stations

Station 123456

s/Ls 0.53 0.59 0.69 0.78 0.88 0.97

x/Cx 0.65 0.72 0.80 0.86 0.92 0.97

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NASA/CR—2012-217415 168
ing from 1.1 mm to over 40 mm.

Flow direction in a separation bubble cannot be determined

with a single-sensor hot-wire, but velocity magnitude can be measured and was found to be near zero within the bubbles of the

present cases when the flow was laminar. In cases where the flow

became turbulent but remained separated, fluctuating velocities

caused false high mean velocity readings in the separation bubble.

With the exception of these turbulent separated cases, the uncertainty in the mean velocity is 3–5% except in the very near-wall

region, where near-wall corrections 32 were applied to the mean

velocity. Uncertainties in the momentum and displacement thicknesses computed from the mean profiles are 10%. The uncertainty

in the shape factor, H, is 8%. Local skin friction coefficients were

computed from the near-wall mean velocity profiles using the

technique of Volino and Simon 33. This technique accounts for

streamwise pressure gradient effects on the mean profile. The uncertainty in Cf is 8%. The uncertainty in the fluctuating streamwise velocity is below 10%.

Pressure surveys and velocity profiles were acquired at nominal

Re= 25,000, 50,000, 100,000, 200,000, 300,000, and 330,000. An

additional pressure survey was acquired at Re= 125,000. The Reynolds number, as defined above, is based on the suction surface

length and the nominal cascade exit velocity. The corresponding

Reynolds numbers based on the cascade inlet velocity and the

axial chord length ranged from 10,000 to 133,000.

3 Results

3.1 Pressure Profiles. Pressure profiles for all cases are

shown in Fig. 4. The inviscid profile for the L1A airfoil is shown

for comparison. At the three highest Re, with the exception of a

slightly higher suction side peak, the data generally agree with the

inviscid profile. This indicates that the boundary layer is attached

over most of the airfoil in these cases. Some deviation between

the three highest Re cases is visible at s/Ls= 0.6 with the Cp value

rising slightly at the lower Re. This may indicate a small separation bubble at this location. At the three lowest Reynolds numbers,

the suction peak is suppressed and the Cp values have a constant

value on the downstream part of the suction side. This plateau

indicates that the boundary layer has separated. The Cp values

never return to the inviscid line, indicating that the boundary layer

never reattaches i.e., the separation bubble bursts. At Reynolds

numbers of 100,000 and below, the separation bubble always appeared to burst. At Re above 150,000 the boundary layer was

attached over most of the airfoil. Between Re= 100,000 and

150,000 the behavior was less predictable, with some dependence

on whether the desired wind tunnel velocity was approached from

above or below. The Re= 124,500 case in Fig. 4 shows the suppressed suction peak and burst bubble of the lower Re cases, but

in some trials not shown the Cp values dropped somewhat from

their plateau near the trailing edge.

The pressure side Cp values also show some change with Reynolds number. Near the leading edge, Cp is higher for the lower

Re cases, suggesting more of a leading edge separation bubble.

Near the trailing edge the Cp values are again higher at the lower

Re. This is believed to result when the suction side separation

bubble forces fluid toward the pressure side, thereby increasing

the pressure side velocity.

The lift on the airfoil can be determined by integrating the

difference between the suction and pressure side Cp values along

the axial direction. The result is shown as a function of Reynolds

number in Fig. 5. The lift is about 20% lower for the separated

flow cases.

The results in Fig. 4 contrast with the results for the Pack B

airfoil presented by Volino 13, Bons et al. 4, Simon et al. 34,

and others. With low freestream turbulence, at Re50,000 the

boundary layer on the Pack B airfoil separated and did not reattach, much like the L1A behavior of the present study. At Re

300,000, the boundary layer remained attached over most of the

surface, again similar to the L1A behavior shown above. At intermediate Reynolds numbers, however, the Pack B results showed

separation with reattachment after a large separation bubble. At

Re= 100,000, for example, the boundary layer on the Pack B airfoil separated just after the suction peak and did not reattach until

near the trailing edge. As Reynolds number increased, the reattachment location gradually moved upstream. The L1A does not

show this large separation bubble behavior, instead appearing to

switch more abruptly between a nearly fully attached boundary

layer and a burst bubble. This will be explored in more detail with

the velocity profiles below.

3.2 Total Pressure Losses. The loss coefficient, 	, is shown

for all cases in Fig. 6. The coordinate  indicates the distance in

the direction perpendicular to the axial chord. The normalizing

quantity L is the blade spacing. The origin, = 0, corresponds to

the location directly downstream of the trailing edge of the center

blade in the direction of the exit design flow angle. At the two

highest Reynolds numbers, the losses are low, and the locations of

the loss peaks are in the expected positions downstream of the

airfoils. This indicates that the actual flow angle is about equal to

the design angle. The peaks downstream of blades B3–B5 are

similar to each other, indicating periodicity in the cascade. The

loss coefficient is near zero between the peaks. At Re= 205,900,

the loss coefficients are slightly larger than at the higher Re. Although the difference is within the uncertainty, the trend is consistent with the appearance of a small separation bubble at this

Reynolds number, seen in Fig. 4 at s/Ls= 0.6. At the lower Reynolds numbers, the burst separation bubble results in much higher

losses, and forces the peaks about 0.35L toward the pressure side

of each passage. This shift corresponds to a 14 deg change in the

exit flow angle. The reduction in flow turning is consistent with

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

Re=25,800

Re=50,900

Re=102,000

Re=124,500

Re=205,900

Re=297,300

Re=328,300

Inviscid

Fig. 4 Cp profiles

0 0.5 1 1.5 2 2.5 3 3.5

x 10

5

0.9

1

1.1

1.2

1.3

Re

∫(Cpsuction−Cppressure)dx/Cx

Fig. 5 Lift based on integrated Cp profile

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NASA/CR—2012-217415 169
the lower lift shown in Fig. 5. The peaks become noticeably

smaller moving from B6 to B3, indicating the effect of the tailboard in reducing the separation bubble thickness.

The integrated loss around the center blade is computed as

	int = 

−L/2

L/2

	Ud/

−L/2

L/2

Ud 2

and is shown in Fig. 7 as a function of Reynolds number. Losses

increase with decreasing Re. Between Re= 300,000 and 200,000,

	int increases by 37%, due to the thicker boundary layer at lower

Reynolds numbers. The most dramatic increase, however, occurs

between the attached and separated flow cases. The integrated loss

increases by a factor of 7 between the highest and lowest Reynolds numbers.

The turbulence associated with the airfoil wakes is shown in

Fig. 8. The rms fluctuating streamwise velocity at 0.63Cx downstream of the trailing edge is normalized on the nominal exit velocity for each case. At the higher Reynolds numbers, the boundary layer turbulence has decayed to about 4% of Ue. A double

peak is apparent, with the slightly higher peak resulting from the

suction side boundary layer and the smaller peak from the pressure side. Between blades the turbulence intensity is at the background level of the wind tunnel. At the lower Reynolds numbers,

the turbulence level is much higher and in agreement with the loss

coefficients, the peaks are shifted toward the pressure side of the

passages.

3.3 Velocity Profiles. Velocity profiles for the six suction surface measurement stations are shown in Fig. 9 for the nominal

Re= 25,000 case. The top row in the figure shows the distance

from the wall normalized on the suction surface length plotted

against the local mean velocity normalized on the nominal exit

velocity, Ue. The boundary layer has just separated at the first

measurement station and the separation bubble grows larger at the

downstream stations. The boundary layer does not reattach. The

second row in Fig. 9 shows the rms streamwise fluctuating velocity, u, normalized with Ue. There is a very large peak located in

the shear layer over the separation bubble, which reaches a dimensionless magnitude of about 0.2. The third row in Fig. 9 shows the

intermittency, 
. The intermittency is the fraction of time the flow

is turbulent. It was determined at each measurement location

based on the instantaneous streamwise velocity signal, using the

technique described in Ref. 35. Turbulent flow is defined here to

include a range of large and small scale eddies. A boundary layer

or shear layer may have significant ufluctuations but still be

considered nonturbulent if these fluctuations are induced by an

external source such as freestream turbulence or are associated

with instability in a narrow frequency band. Transition to turbulence is characterized by the appearance of broadband fluctuations. In the intermittency processing routine, the velocity signal

is high pass filtered, and the appearance of high frequency fluctuations is used to distinguish between turbulent and nonturbulent

flows. The results in Fig. 9 indicate that the shear layer remains

nonturbulent until the last measurement station, where a small

peak indicates the possible beginning of transition.

Figure 10 presents another view of the transition process using

the turbulence spectra. The spectra are computed from the fluctuating velocity signal acquired at the location of peak uat each

−1.5 −1 −0.5 0 0.5 1 1.5 0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

φ/Lφ

ψ

B3

B4 B5

B6

Re=25,800

Re=50,900

Re=102,000

Re=124,500

Re=205,900

Re=297,300

Re=333,900

Fig. 6 Total pressure loss coefficient at 0.63Cx downstream of

cascade

0 0.5 1 1.5 2 2.5 3 3.5

x 105

0

0.1

0.2

0.3

0.4

0.5

Re

ψint

Fig. 7 Integrated total pressure loss coefficient for center

blade as function of Re

−1.5 −1 −0.5 0 0.5 1 1.5 0

0.05

0.1

0.15

0.2

0.25

φ/Lφ

u′/Ue

B3 B4 B5

B3

B4

B5

B6

Re=25,800

Re=50,900

Re=102,000

Re=124,500

Re=205,900

Re=297,300

Re=333,900

Fig. 8 rms fluctuating streamwise velocity at 0.63Cx downstream of cascade

0 0 0 0 0 0 1 0

0.1

0.2

U/Ue

y/Ls

a) 1 2 3 4 5 6

0 0 0 0 0 0 0.1 0

0.1

0.2

u′/Ue

y/Ls

b) 1 2 3 4 5 6

0 0 0 0 0 0 1 0

0.1

0.2

γ

y/Ls

c) 1 2 3 4 5 6

Fig. 9 Profiles for Re=25,000 case: „a... mean velocity, „b...

u/Ue, and „c... intermittency

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NASA/CR—2012-217415 170
measurement station shown in Fig. 9. Frequencies are resolved

from 4.88 Hz to 10 kHz in 4.88 Hz increments using 4096 point

fast Fourier transforms to compute the spectra. The frequencies

and power spectral density PSD are normalized using Ls and Ue.

There is a clear broadband rise in the spectra between the third

and fourth measurement stations, and the magnitude continues to

rise downstream. The rise indicates that transition may be beginning by the fourth station. The rise in the spectra is not as abrupt

as at higher Reynolds numbers shown below, however, which

may indicate that transition is only in its earliest stages. This may

explain the intermittency values near zero in Fig. 9.

The velocity profiles for the Re= 50,000 case are shown in Fig.

11. The mean and fluctuating velocity show essentially the same

behavior observed in the Re= 25,000 case of Fig. 9. The boundary

layer does not reattach, in agreement with the pressure profile of

Fig. 4. The intermittency values rise slightly above zero at the

third station and indicate that transition is clearly underway by the

fourth station. The corresponding turbulence spectra of Fig. 12

agree, showing a clear rise in the power spectrum between the

second and third stations and a further rise to a more turbulent

state by the fourth station.

The initiation of transition without boundary layer reattachment

shown in Figs. 11 and 12 is markedly different than the behavior

observed in previous studies with other LPT airfoils. Volino 13

noted that on the Pack B airfoil, when transition began in the

separated shear layer, the associated mixing almost immediately

induced boundary layer reattachment. This behavior was utilized

by Volino 22 and by Zhang et al. 20, who used small passive

devices on their airfoils to control separation. These devices were

too small to trip the boundary layer to turbulent but introduced

small disturbances, which accelerated the transition process in the

shear layer, thereby moving reattachment upstream. Both Volino

22 and Zhang and Hodson 27 found that these small devices

resulted in lower losses than large devices, which immediately

tripped the boundary layer to turbulent. The present results suggest that such devices may not work with the L1A airfoil because

transition is not sufficient to force reattachment. The strong negative acceleration parameter, shown in Fig. 2, particularly at low

Re is apparently strong enough to prevent reattachment of the

turbulent shear layer. A comparison of the present results to those

of Volino 13 shows that the stronger pressure gradient results in

a separation bubble on the L1A that is about four times thicker

than that on the Pack B airfoil in terms of y /Ls at a given Reynolds number. When transition occurs, the turbulence in the shear

layer is apparently too far from the airfoil to induce enough nearwall mixing to force reattachment. This suggests that transition,

whether naturally occurring or induced through flow control, must

occur far enough upstream to cause reattachment before the separation bubble becomes too thick.

Figure 13 shows the velocity profiles for the Re= 100,000 case.

Similar to the previous cases, the separation bubble grows in the

streamwise direction, the boundary layer does not reattach, and

there is a very high peak in the fluctuating velocity in the shear

layer over the separation bubble. The intermittency rises slightly

above zero at the second station and indicates fully turbulent flow

by the third station, where 
= 1 in the shear layer. As in the

10−1 100 101 102

10−10

10−8

10−6

10−4

10−2

100

f⋅Ls/Ue

PSD/(Ls

⋅Ue

)

s/Ls=0.531

0.594

0.688

0.781

0.875

0.969

Fig. 10 Turbulence spectra for Re=25,000 case

0 0 0 0 0 0 1 0

0.1

0.2

U/Ue

y/Ls

a) 1 2 3 4 5 6

0 0 0 0 0 0 0.1 0

0.1

0.2

u′/Ue

y/Ls

b) 1 2 3 4 5 6

0 0 0 0 0 0 1 0

0.1

0.2

γ

y/Ls

c) 1 2 3 4 5 6

Fig. 11 Profiles for Re=50,000 case: „a... mean velocity, „b...

u/Ue, and „c... intermittency

10−1 100 101 102

10−10

10−8

10−6

10−4

10−2

100

f⋅Ls/Ue

PSD/(Ls

⋅Ue

)

s/Ls=0.531

0.594

0.688

0.781

0.875

0.969

Fig. 12 Turbulence spectra for Re=50,000 case

0 0 0 0 0 0 1 0

0.1

0.2

U/Ue

y/Ls

a) 1 2 3 4 5 6

0 0 0 0 0 0 0.1 0

0.1

0.2

u′/Ue

y/Ls

b) 1 2 3 4 5 6

0 0 0 0 0 0 1 0

0.1

0.2

γ

y/Ls

c) 1 2 3 4 5 6

Fig. 13 Profiles for Re=100,000 case: „a... mean velocity, „b...

u/Ue, and „c... intermittency

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NASA/CR—2012-217415 171
previous case, the transition is not sufficient to induce boundary

layer reattachment. Figure 14 shows the spectra for the Re

= 100,000 case. There is a clear narrowband rise in the spectrum

between the first and second stations, centered about a dimensionless frequency of 12. A further broadband rise occurs by the third

station, indicating a turbulent shear layer, in agreement with the

intermittency profiles of Fig. 13. The peak at the second station in

Fig. 14 indicates a shear layer instability, which likely initiates

transition. Similar peaks are visible in the other Reynolds number

cases, and will be discussed further below.

Velocity profiles for the Re= 200,000 case are shown in Fig. 15.

The boundary layer was laminar and on the verge of separation at

the first two measurement stations. The small peak in intermittency at the first measurement station is believed to be erroneous,

based on the spectra and the zero intermittency measurement at

the second station. The skin friction, as determined from the nearwall velocity profile, was near zero at the first two stations, but

any separation bubble was very thin. Between the second and

third stations, the boundary layer underwent transition. The

boundary layer was clearly attached and fully turbulent at the third

station and remained attached at all downstream stations, in agreement with the pressure profiles of Fig. 4. The spectra for this case

are shown in Fig. 16. The clear jump from laminar to turbulent

flow is clear between the second and third stations.

The velocity profiles and spectra for the Re= 300,000 case are

shown in Figs. 17 and 18. The results are very similar to those of

the Re= 200,000 case, with transition occurring between the second and third stations, and the boundary layer remaining attached

and turbulent downstream.

The shape factor, H, and skin friction coefficient Cf are good

indicators of the state of the boundary layer with respect to separation and transition. The shape factor is shown for the high Re

cases in Fig. 19. It is between 2.5 and 3.5 at the first two measurement stations, where the boundary layer is laminar and on the

verge of separation. It then drops rapidly with the onset of transi10−1 100 101 102

10−10

10−8

10−6

10−4

10−2

100

f⋅Ls/Ue

PSD/(Ls

⋅Ue

)

s/Ls=0.531

0.594

0.688

0.781

0.875

0.969

Fig. 14 Turbulence spectra for Re=100,000 case

0 0 0 0 0 0 1 0

0.02

0.04

0.06

0.08

U/Ue

y/Ls

a) 1 2 3 4 5 6

0 0 0 0 0 0 0.1 0

0.02

0.04

0.06

0.08

u′/Ue

y/Ls

b) 1 2 3 4 5 6

0 0 0 0 0 0 1 0

0.02

0.04

0.06

0.08

γ

y L/ s

c) 1 2 3 4 5 6

Fig. 15 Profiles for Re=200,000 case: „a... mean velocity, „b...

u/Ue, and „c... intermittency

10−1 100 101 102

10−10

10−8

10−6

10−4

10−2

100

f⋅Ls/Ue

PSD/(Ls

⋅Ue

)

s/Ls=0.531

0.594

0.688

0.781

0.875

0.969

Fig. 16 Turbulence spectra for Re=200,000 case

0 0 0 0 0 0 1 0

0.02

0.04

0.06

0.08

U/Ue

y L/ s

a) 1 2 3 4 5 6

0 0 0 0 0 0 0.1 0

0.02

0.04

0.06

0.08

u′/Ue

y/Ls

b) 1 2 3 4 5 6

0 0 0 0 0 0 1 0

0.02

0.04

0.06

0.08

γ

y L/ s

c) 1 2 3 4 5 6

Fig. 17 Profiles for Re=300,000 case: „a... mean velocity, „b...

u/Ue, and „c... intermittency

10−1 100 101 102

10−10

10−8

10−6

10−4

10−2

100

f⋅Ls/Ue

PSD/(Ls

⋅Ue

)

s/Ls=0.531

0.594

0.688

0.781

0.875

0.969

Fig. 18 Turbulence spectra for Re=300,000 case

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NASA/CR—2012-217415 172
tion, reaching a minimum of about 1.57 at the fourth measurement

station. While this value clearly indicates that the boundary layer

is attached and turbulent, it is still above the zero pressure gradient turbulent value of about 1.4. At the downstream stations, H

begins to rise, and it exceeds 1.7 by the last station. The skin

friction coefficient is shown as a function of momentum thickness

Reynolds number in Fig. 20. It rises from near zero after transition

and decreases in the streamwise direction. Also shown is a standard flat-plate correlation for turbulent boundary layers from

Schlichting 36 and the Ludwieg–Tillmann correlation. The

present Cf values are as much as 40% lower than the flat-plate

correlation and agree well with the Ludwieg–Tillmann correlation. The rising H and low Cf values are indicative of the strong

adverse pressure gradient. The corresponding values on the Pack

B airfoil 13 were closer to typical zero pressure gradient values,

which is consistent with the weaker pressure gradient on the Pack

B and the stronger tendency for reattachment at low Re on the

Pack B.

3.4 Instability Prediction. The spectral peak noted above at

the second station in Fig. 14 suggests an instability in the shear

layer that likely initiates transition. Although less distinct, similar

peaks are visible just before or just after transition inception at all

Reynolds numbers. Volino 14 observed similar peaks in the

boundary layer on the Pack B airfoil and associated them with TS

waves, using an analysis from Ref. 37. The TS waves are believed to form in the attached boundary layer between the beginning of the adverse pressure gradient region and the separation

location. The TS waves then continued to grow in the separated

shear layer until they became large enough to induce transition, as

discussed in Ref. 18. Walker 37 gave the most unstable frequency for TS waves as

2f/U


2 = 3.2 Re

−3/2 3

Hughes and Walker 31 noted that the frequency predicted by Eq.

3 is a function of streamwise position, since the freestream velocity, U
, varies in nonzero pressure gradient cases, and the displacement thickness, , changes as the boundary layer grows.

Hence, a single frequency cannot be expected for all TS waves in

a boundary layer. For the present study, U
 and Rewere taken at

the separation location. The Revalue was estimated with the

boundary layer code TEXSTAN 38. The code was used to compute

the laminar suction side boundary layer from the leading edge to

the separation point using the pressure gradient from Fig. 4 as

input for each case. The code stops when the flow separates, and

the separation points computed were in good agreement with the

pressure and velocity profiles presented above. Figure 21 shows

the peak frequencies extracted from Figs. 10, 12, 14, 16, and 18

along with frequencies predicted by Eq. 3. The good agreement

between the experimental and predicted values indicates that TS

waves may play a role in the transition. It is also possible that

other instability mechanisms are present. Stieger and Hodson 39,

and Roberts and Yaras 40, for example, observed Kelvin–

Helmholtz rollup of the shear layer.

3.5 Transition Correlations. Several correlations for predicting the starting location for separated flow transition are available

in the literature, and some of these are tested below against the

present data.

Mayle 2 presented the following correlations for short and

long separation bubbles.

Rest = 300 Res

0.7 short bubble 4

Rest = 1000 Res

0.7 long bubble 5

where Rest is the Reynolds number based on the freestream velocity at separation and the distance from separation to transition

start.

The correlation of Davis et al. 16 is

Rest = 25000 log10coth0.1732TI 6

For the present experiments, TI could reasonably be set between

0.3% and 0.8%, as discussed above. Varying TI within this range

does not significantly affect the correlation predictions. For Eq.

6 and all the correlations below, TI was set to 0.5%.

Hatman and Wang 15 identified several transition modes and

present correlations for each of them. Their laminar separation

mode transition correlation can be cast in terms of Rest as

Rest = 0.0816 Res + 26805 7

Yaras 17 proposed

Rest = 0.04 Res + 6.3  104

1 − tanh3TF 8

where TF=maxTF, 1%, and TF=TIss/

0.2, where ss is the

distance from the leading edge to the separation point and  is the

integral length scale of the freestream turbulence.

Roberts and Yaras 40 presented

Rest = 785 − 30TFRes

0.7 9

Praisner and Clark 19 presented the correlation

Rest = 173.0 Res Res

−1.227 10

The above correlations are based on the conditions at the separation location and in some cases the freestream turbulence inten0.5 0.6 0.7 0.8 0.9 1 1.5

2

2.5

3

3.5

s/Ls

H

Re=205,900

Re=297,300

Re=333,900

Fig. 19 Shape factor for high Re cases

200 400 600 800 1000 1200 1400 0

2

4

6

8

x 10−3

Cf=0.0256/Reθ

0.25

Cf=0.246⋅ 10−0.678H Reθ

−0.268

Reθ

Cf

Re=205,900

Re=297,300

Re=333,900

Flat plate

Ludwieg−Tillmann

Fig. 20 Skin friction coefficient for high Re cases

200 300 400 500 600 700 800 900 0

500

1000

1500

2000

Reδ*

f [Hz]

Experiment

Walker [37] correlation

Fig. 21 Frequencies of spectral peaks along with predicted

most unstable TS frequencies at separation location

011007-8 / Vol. 132, JANUARY 2010 Transactions of the ASME

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NASA/CR—2012-217415 173
sity. Volino and Bohl 18 reasoned that instabilities begin to grow

when the boundary layer becomes unstable at the start of the

adverse pressure gradient region, and presented the following correlation:

Rept = 8.806.37 − log10TI2

Rep

4/3 11

where Rept is based on the freestream velocity at the suction peak

and the distance from the suction peak to the transition start.

The location of the suction peak for each case was taken from

the experimental data of Fig. 4. The separation location and the

momentum thickness Reynolds numbers at separation, Res, and

at the suction peak, Rep, were taken from the TEXSTAN calculations described above. Equations 4–11 were then used to compute st, which is defined as the distance from the leading edge to

the start of transition. The Reynolds numbers used with the correlations are listed in Table 3. The st results are shown as a function of Re in Fig. 22. The experimental transition locations, as

determined from Figs. 9–18, are also shown. The uncertainties in

the experimental locations, which result from the finite spacing

between the measurement stations, are indicated by the error bars

in Fig. 22. Values of st /Ls1 in the figure indicate that transition

is predicted downstream of the trailing edge of the airfoil.

The Mayle 2 short bubble correlation, which is only intended

for separation bubbles that reattach, predicts transition somewhat

too far upstream at all Reynolds numbers. For the Re200,000

cases, all of the other correlations agree with the experimental

results. For the lower Re cases, the Roberts and Yaras 40 and

Volino and Bohl 18 correlations continue to agree with the experiment, while the other correlations predict transition too far

downstream.

4 Conclusions

The flow over the very high lift L1A airfoil was studied experimentally using a linear cascade. Reynolds numbers based on the

suction surface length and nominal exit velocity ranging from

25,000 to 330,000 were considered. In all cases the laminar suction surface boundary layer separated, but at Reynolds numbers

greater than 150,000, the separation bubble was very thin and

short, and the boundary layer was attached over most of the surface. At lower Reynolds numbers, the boundary layer separated

and never reattached. Separation without reattachment caused the

lift on the airfoil to decrease by 20% and the total pressure losses

increased by a factor of 7 above the highest Reynolds number

case. Transition to turbulence occurred in all cases in the shear

layer after separation, and appeared to be caused by the growth of

Tollmien–Schlichting waves, which originated in the boundary

layer upstream of separation. The transition location was well predicted by the correlations of Roberts and Yaras 40 and Volino

and Bohl 18. Transition caused immediate reattachment in the

high Reynolds number cases, but the turbulent shear layer remained separated in the low Re cases. This behavior contrasts

with previous studies on other LPT airfoils. In those studies, transition immediately triggered reattachment, even at low Reynolds

numbers. On the present airfoil, the strong adverse pressure gradient prevents reattachment at low Re, even after transition occurs.

Acknowledgment

This work was sponsored by the National Aeronautics and

Space Administration Award No. NNC07IA10I. The grant monitor is Dr. Anthony Strazisar of the NASA Glenn Research Center.

The support of the United States Naval Academy Technical Support Department Shop and Fluids Laboratory is greatly appreciated.

Nomenclature

Cf  skin friction coefficient

Cp  2PT−P/Ue

2

, pressure coefficient

Cx  axial chord length

f  frequency

H  /, shape factor

K  /U


2 dU
/ds, acceleration parameter

Ls  suction surface length

L  blade spacing pitch

P  pressure

PS  upstream static pressure

PT  upstream stagnation pressure

PTe  downstream stagnation pressure

Re  UeLs/, exit Reynolds number

Re displacement thickness Reynolds number

Re  momentum thickness Reynolds number

s  streamwise coordinate, distance from leading

edge

TI  freestream turbulence intensity

U  mean streamwise velocity

U
  local freestream velocity

Ue  nominal exit freestream velocity, based on inviscid solution

u time averaged rms streamwise fluctuating

velocity

x  axial distance from leading edge

y  distance from wall

Zw  Zweifel coefficient

1  inlet flow angle

2  exit flow angle

 displacement thickness

  coordinate along blade spacing, normal to axial

chord

Table 3 Conditions at suction peak and separation location

based on laminar boundary layer calculation

Re sp /Ls ss/Ls Rep Res

25,800 0.438 0.496 48 61

50,900 0.438 0.494 68 85

102,000 0.438 0.496 96 122

205,900 0.493 0.538 151 182

297,300 0.493 0.538 182 220

333,900 0.493 0.537 193 231

0 0.5 1 1.5 2 2.5 3 3.5

x 10

5

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

Re

st/Ls

Volino and Bohl [18]

Mayle [2] short bubble

Mayle [2] long bubble

Davis et al. [16]

Hatman and Wang [15]

Yaras [17]

Roberts and Yaras [40]

Praisner and Clark [19]

Experiment

Fig. 22 Predicted and experimental transition start location:

error bars indicate uncertainty due to finite spacing of measurement stations

Journal of Turbomachinery JANUARY 2010, Vol. 132 / 011007-9

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NASA/CR—2012-217415 174

  intermittency, fraction of time flow is turbulent

  kinematic viscosity

  density

  momentum thickness

	  PT−PTe/PT−PS, total pressure loss

coefficient

	int  total pressure loss integrated over blade

spacing

Subscripts

p  pressure minimum suction peak location

pt  distance from suction peak to transition start

s  separation location

st  distance from separation location to transition

start


  freestream

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011007-10 / Vol. 132, JANUARY 2010 Transactions of the ASME

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NASA/CR—2012-217415 175
Ralph J. Volino

Department of Mechanical Engineering,

United States Naval Academy,

Annapolis, MD 21402-5042

e-mail: volino@usna.edu

Olga Kartuzova

e-mail: kartuzova_olga@hotmail.com

Mounir B. Ibrahim

e-mail: m.ibrahim@csuohio.edu

Department of Mechanical Engineering,

Cleveland State University,

Cleveland, OH 44115-2425

Separation Control on a Very High

Lift Low Pressure Turbine Airfoil

Using Pulsed Vortex Generator

Jets

Boundary layer separation control has been studied using vortex generator jets (VGJs)

on a very high lift, low-pressure turbine airfoil. Experiments were done under high (4%)

freestream turbulence conditions on a linear cascade in a low speed wind tunnel. Pressure surveys on the airfoil surface and downstream total pressure loss surveys were

documented. Instantaneous velocity profile measurements were acquired in the suction

surface boundary layer. Cases were considered at Reynolds numbers (based on the suction surface length and the nominal exit velocity from the cascade) of 25,000 and 50,000.

Jet pulsing frequency, duty cycle, and blowing ratio were all varied. Computational

results from a large eddy simulation of one case showed reattachment in agreement with

the experiment. In cases without flow control, the boundary layer separated and did not

reattach. With the VGJs, separation control was possible even at the lowest Reynolds

number. Pulsed VGJs were more effective than steady jets. At sufficiently high pulsing

frequencies, separation control was possible even with low jet velocities and low duty

cycles. At lower frequencies, higher jet velocity was required, particularly at low Reynolds numbers. Effective separation control resulted in an increase in lift and a reduction

in total pressure losses. Phase averaged velocity profiles and wavelet spectra of the

velocity show the VGJ disturbance causes the boundary layer to reattach, but that it can

reseparate between disturbances. When the disturbances occur at high enough frequency,

the time available for separation is reduced, and the separation bubble remains closed at

all times. DOI: 10.1115/1.4003024

1 Introduction

Boundary layer separation can lead to partial loss of lift and

higher aerodynamic losses on low-pressure turbine LPT airfoils

e.g., Refs. 1–3. As designers impose higher loading to improve

efficiency and lower cost, the associated strong adverse pressure

gradients on the suction side of the airfoil can exacerbate separation problems. In aircraft engines, the lower density and therefore

lower Reynolds numbers at altitude can lead to a component efficiency drop of 2% between takeoff and cruise in large commercial transport engines, and possibly as much as 7% in smaller

engines operating at higher altitudes 4,5. Prediction and control

of suction side separation, without sacrifice of the benefits of

higher loading, are, therefore, crucial for improved engine design.

Separation and separated flow transition, which can lead to

boundary layer reattachment, have received considerable attention. Lists of various studies are provided by Volino 6,7. In

general, previous work shows that the strong acceleration on the

leading section of the airfoil keeps the boundary layer thin and

laminar, even in the presence of elevated freestream turbulence.

When separation does occur, it is usually just downstream of the

suction peak. If transition then occurs in the shear layer over the

separation bubble, it is typically rapid and often causes the boundary layer to reattach 7,8.

Airfoils can be designed with high resistance to separation, as

described by Praisner and Clark 9, but a loading limit will always exist, above which separation will still occur. If flow control

were incorporated in the design of an advanced airfoil, as discussed by Bons et al. 10, it might be possible to increase the

loading limit. Separation control with passive devices such as

boundary layer trips has been shown effective by Zhang et al.

11, Bohl and Volino 12, Volino 13, and others. Passive devices have the distinct advantage of simplicity, but they also introduce parasitic losses and cannot be adjusted to account for

changes in flow conditions. Active devices could potentially provide better control over the entire operating range of interest and

be reduced in strength or turned off to avoid unnecessary losses

when they are not needed. Potential disadvantages of active devices are the cost of implementation and the possibility of control

failure, particularly in aircraft engines.

The literature contains many examples of active separation control. A few which could be applied in turbomachinery are discussed in Ref. 14. Plasma devices, as used by Huang et al. 15,

could be viable, and are under active study. Vortex generator jets

VGJs, as introduced by Johnston and Nishi 16, are another

alternative and the subject of the present study. Blowing from

small, compound angled holes is used to create streamwise vortices. The vortices bring high momentum fluid into the near wall

region, which can help to control separation. The most effective

VGJs enter the boundary layer at a relatively shallow pitch angle

typically 30–45 deg relative to the wall and a high skew angle

45–90 deg relative to the main flow. Additionally, the jets promote transition, and turbulent mixing also helps to mitigate separation. Bons et al. 10noted that in the case of pulsed VGJs, the

turbulence effect is more significant than the action of the vortices. Bons et al. 4,17, Volino 14, Volino and Bohl 18, McQuilling and Jacob 19, and Eldredge and Bons 20all used

VGJs on the highly loaded Pack B LPT airfoil. Separation was

essentially eliminated, even at the lowest Reynolds number considered Re=25,000 based on suction surface length and nominal

exit velocity. Pulsed jets were more effective than steady jets.

The initial disturbance created by each pulse caused the boundary

Contributed by the International Gas Turbine Institute IGTI of ASME for publication in the JOURNAL OF TURBOMACHINERY. Manuscript received June 29, 2010; final

manuscript July 11, 2010; published online April 25, 2011. Editor: David Wisler.

Journal of Turbomachinery Copyright © 2011 by ASME OCTOBER 2011, Vol. 133 / 041021-1

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NASA/CR—2012-217415 176
layer to attach. The turbulence was followed by a calmed period

21,22during which the boundary layer was very resistant to

separation, much like a turbulent boundary layer, but very laminarlike in terms of its fluctuation levels and low losses. When the

time between pulses was long enough, the boundary layer did

eventually relax to a separated state, but due to the control which

persisted during the calmed period, the VGJs were effective even

with low jet pulsing frequencies, duty cycles, and mass flow rates.

Since the boundary layer was attached and undisturbed for much

of the jet pulsing cycle, profile losses were low.

Similar results with pulsed VGJs were found on the L1M airfoil

by Bons et al. 23. The L1M is more highly loaded than the Pack

B, but more resistant to separation because of forward loading. A

large separation bubble followed by boundary layer reattachment

was observed at low Reynolds numbers, and pulsed VGJs reduced

the size of the bubble.

In the present study, the very highly loaded L1A airfoil is used.

The L1A was designed at the Air Force Research Laboratory

AFRL and is available on a limited basis from Ref. 24. It is an

aft loaded blade with the same flow angles and loading as the

L1M. Dimensions of the L1A as used in the present study are

given in Table 1. Based on the design calculations of Clark 24,

the L1A has a Zweifel coefficient of 1.35, which corresponds to

10% higher loading than the “ultrahigh lift” airfoils described by

Zhang and Hodson 25, and 17% higher loading than the Pack B.

Because the L1A is aft loaded, it is more prone to separation than

the L1M, as documented in Refs. 26,6,27,28. In cases without

flow control and low Reynolds numbers, the boundary layer separates and does not reattach, in spite of transition to turbulence in

the shear layer over the separation bubble. This result contrasts

with the results of studies on less aggressive airfoils, which all

showed reattachment after transition. The separation bubble on the

L1A is about four times thicker than that on the Pack B. The

larger distance from the shear layer to the wall on the L1A apparently prevents the turbulent mixing in the shear layer from reaching the wall and causing reattachment. The failure of the boundary

layer to reattach results in a 20% loss in lift and increases profile

losses by up to a factor of 7. At higher Reynolds numbers, the

separation bubble is small and the boundary layer is attached over

most of the airfoil.

Separation control with VGJs has been demonstrated on the

L1A airfoil. Bons et al. 26considered a case with Re=50,000,

background freestream turbulence TI=3%, and periodic wakes

produced with moving rods upstream of the airfoils. The VGJs

had a duty cycle fraction of time the jets are on during a pulsing

cycle D=30% and a blowing ratio B=2.3 based on the maximum jet velocity in the pulsing cycle and the local freestream

velocity. The dimensionless frequency of the wake passing and

VGJ pulsing were both F= fLj−te /Uave=0.34, where Lj−te is the

streamwise distance from the VGJ holes to the trailing edge, and

Uave is the average freestream velocity over this distance. Two

different streamwise locations for the VGJs were considered. The

VGJs reduced the separation bubble size and reduced total pressure losses. Volino et al. 29considered cases with TI=0.6%.

Pressure distributions on the airfoils and total pressure losses were

documented. Cases were considered at Reynolds numbers from

25,000 to 100,000. Jet pulsing frequency was varied from F

=0.14 to 1.12 with duty cycles of 10% and 50% and blowing

ratios ranging from 0.25 to 3.0. In agreement with previous studies, pulsed jets were more effective than steady jets. At Re

=25,000, separation control was achieved with B=0.75, F=0.56,

and D=10%. Without flow control the lift at Re=25,000 was

about 80% of the high Re lift, and with flow control it was 94% of

the high Re lift. Flow control caused the total pressure loss to drop

from about five times the high Re value to about three times the

high Re value. Once good control was achieved, further increasing the frequency, blowing ratio, or duty cycle provided little additional benefit. Partial control was possible at lower frequencies,

particularly with higher blowing ratios or duty cycles. At Re

=50,000 with F=0.56, B=0.25 was sufficient with D=10%.

Lower frequencies were more effective at Re=50,000 than at

Re=25,000. Effective separation control resulted in a 20% increase in lift, and up to a 60% reduction in total pressure loss,

dropping from about five times the high Re value to about twice

the high Re value.

In the present study, flow control with vortex generator jets is

investigated under high 4% freestream turbulence conditions,

which is more representative of engine conditions than the low TI

considered in Ref. 29. Cases with a range of pulsing frequencies

and blowing ratios are considered. In addition to pressure distributions, instantaneous boundary layer velocity measurements are

used to explain the mechanism by which the VGJs control separation.

2 Experimental Facility and Measurements

Experiments were conducted in a closed loop wind tunnel with

a seven blade linear cascade located in the wind tunnel’s third

turn, as shown in Fig. 1a. A coarse grid, consisting of a 1.5 mm

thick sheet metal plate with 19 mm square holes spaced 25.4 mm

apart, center to center, in both directions is located upstream of the

cascade. In a plane perpendicular to the inlet flow and 1.7Cx upstream of the center blade, the grid produced uniform flow with

TI=6.0% in the streamwise component and 4.2% in the cross

stream components, for an overall intensity of 4.9%. The integral

length scale of the freestream turbulence is 0.12Cx in the streamwise component and 0.04Cx in the other components. The streamwise component was also measured at the inlet plane of the cascade in the four center passages, where it had decayed to between

4% and 4.2% between blades B2 and B5 and 4.6% between

blades B5 and B6. Downstream of the cascade, the local TI is

1.8% across all passages, as documented in Ref. 28. The local

freestream turbulence intensity in the passage at the beginning of

the adverse pressure gradient region is 1.4%. The change in TI

through the passage is due mainly to the change in the local

freestream velocity along with some decay of the turbulence.

A tailboard, shown in Fig. 1a, is needed to produce the correct

exit flow angle from the cascade. Its position was set to produce

periodicity at high Reynolds numbers. A tailboard on the opposite

side of the cascade, and inlet guide vanes were found to be unnecessary. To produce the correct approach flow to the end blades

B1 and B7, the amount of flow escaping around the two ends of

the cascade is controlled with the flaps shown in Fig. 1a. The

inlet flow angle was checked with a three-hole pressure probe and

found to agree with the design angle to within 2 deg of uncertainty. Good periodicity at high Reynolds numbers was obtained

in the exit flow. At low Reynolds numbers, when significant separation bubbles are present, the periodicity is not as good due to

suppression of the separation bubble thickness on the blades closest to the tailboard. In cases with effective flow control, periodicity is reestablished. The lack of periodicity in cases with large

separation bubbles is considered acceptable since the focus of the

Table 1 Cascade parameters

Axial chord, Cx

mm

True chord

mm

Pitch, L

mm

Span

mm

Suction side, Ls

mm

Inlet flow angle

deg

Exit flow angle

deg

134 146 136 724 203 35 60

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NASA/CR—2012-217415 177
study is separation control and not documentation of cases with

large separation that would be unacceptable in practice. This compromise facilitates the study of a larger number of cases with flow

control by obviating the need to adjust the tailboard by trial and

error for each case. It also provides for better repeatability in the

experiments, since the position of the tailboard is fixed for all

cases. Any improvements made with flow control will be larger in

practice than documented in the experiment, due to the effect of

the tailboard in suppressing the bubble size in the uncontrolled

cases.

Each blade in the cascade has a central cavity, which extends

along the entire span. As explained in Ref. 29, compressed air is

supplied to the cavities from a common manifold. The manifold is

supplied through two fast response solenoid valves Parker Hannifin 009-0339-900 with General Valve Iota One pulse driver

operating in parallel. A single spanwise row of holes was drilled

into the suction surface of each blade at the inviscid pressure

minimum location, s/Ls=0.5 x/Cx=0.62, where s is the distance

from the leading edge and Ls is the suction surface length. The

pressure minimum has been shown in the studies listed above to

be about the optimal location for flow control devices. The holes

are 0.8 mm 0.006Cx in diameter and drilled at 30 deg to the

surface and 90 deg to the main flow direction, as shown in Fig.

1b. This is the same orientation used in all the VGJ studies listed

above. The hole spacing is 10.6 diameters, and the length to diameter ratio is 12. With steady blowing and B=1, the mass flow

rate of the jets is 0.04% of the main flow mass flow rate. With

pulsed jets the mass flow is lower, particularly in low duty cycle

cases. The solenoid valves pulse the VGJs, and the pulsing frequency is presented below in dimensionless form as F

= fLj−te /Uave. For flow over single airfoils, F1 is typically

needed to maintain separation control, but for cascades, Bons et

al. 17showed that control is possible in some cases with F

=0.1. As shown in Refs. 14,17, this is due to the extended

calmed region which follows the jet disturbance. In practice, VGJs

could be timed to wake passing in an LPT, which has a typical

frequency of about F=0.3.

The center blade, designated B4 in Fig. 1, contains pressure

taps near the spanwise centerline. Pressure surveys are made using

a pressure transducer 0–870 Pa range Validyne transducer. Stagnation pressure is measured with a pitot tube upstream of the

cascade. The uncertainty in the suction side pressure coefficients,

Cp, is 0.07. Most of this uncertainty is due to bias error. Stochastic error is minimized by averaging pressure transducer readings

over a 10 s period.

Total pressure losses are documented using a Kiel probe traversed across three blade spacings, 0.63Cx downstream of the

cascade. A traverse is located in the wind tunnel downstream of

the cascade to move the probe. The traverse causes an acceptably

low blockage when it is located at least two Cx downstream of the

cascade.

Pressure and loss surveys were acquired at nominal Re

=25,000 and 50,000. The Reynolds number, as defined above, is

based on the suction surface length and the nominal cascade exit

velocity. The corresponding Reynolds numbers based on the cascade inlet velocity and the axial chord length are 10,000 and

20,000.

The jet velocity was measured at the exit plane of the VGJ

holes and was documented in Ref. 29. There is about a 5 ms

delay between the opening of the solenoid valve and the beginning of the rise of the jet velocity above zero at the exit plane. The

delay corresponds to the time required for a pressure wave to

travel from the valve to the jet hole. The jet velocity rises rapidly

and nearly continuously for another 6 ms to a maximum and then

drops for about 6 ms toward a steady value which is maintained

for the rest of the time the valve is open. When the valve closes,

the jet velocity drops to zero. In cases with high pulsing frequency

or low duty cycle, the jet turns off before a steady velocity is

reached. As will be shown below, the transient that occurs when

the VGJ turns on is more effective for separation control than the

steady blowing which may follow. The initial velocity maximum

is therefore used to define the blowing ratio rather than the average velocity for the pulse.

Velocity profiles on the suction surface were measured at the

six streamwise stations listed in Table 2. All stations are downstream of the inviscid pressure minimum at s/Ls=0.49. Profiles

were acquired near the spanwise centerline of the airfoil with a

hot-wire anemometer AA Lab Systems model AN-1003 and a

single-sensor hot-film probe TSI model 1201-20. The sensor diameter is 51 m, and the active length is 1.02 mm. At each

measurement location, data were acquired for 26 s at a 20 kHz

sampling rate 219 samples. All raw data were saved. The high

sampling rate provides an essentially continuous signal, and the

long sampling time results in low uncertainty in both statistical

and spectral quantities. Data were acquired at 40 wall normal

locations in each profile, extending from the wall to the

freestream, with most points concentrated in the near wall region.

The probe was positioned as close to tangent to the airfoil surface

as possible at each station, such that the probe body extended

downstream of the sensor and the direction of the traverse was

within 5 deg of normal to the surface. In most cases the closest

Fig. 1 Drawings of test section: „a... linear cascade and „b... airfoil with VGJ holes and cross section of hole geometry

Table 2 Velocity profile measurement stations

Station

1 23456

s/Ls 0.53 0.59 0.69 0.78 0.88 0.97

x/Cx 0.65 0.72 0.80 0.86 0.92 0.97

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NASA/CR—2012-217415 178
point to the wall in each profile was within about 0.2 mm of the

wall, which compares to boundary layer thicknesses ranging from

1.1 mm to over 40 mm.

Flow direction in a separation bubble cannot be determined

with a single-sensor hot-wire, but velocity magnitude can be measured and was found to be near zero within the bubbles of the

present cases when the flow was laminar. In cases where the flow

became turbulent but remained separated, fluctuating velocities

caused false high mean velocity readings in the separation bubble.

With the exception of these turbulent separated cases, the uncertainty in the mean velocity is 3–5% except in the very near wall

region, where near wall corrections 30were applied to the mean

velocity.

Velocity data were both time averaged and ensemble averaged

based on the phase within the jet pulsing cycle. Phase averages of

mean and fluctuating velocity are shown below at 24 dimensionless times, t/T, within the pulsing cycle, where t is time and T is

the period of the cycle. The time t/T=0 corresponds to the opening of the solenoid valve.

Wavelet spectra of the fluctuating velocity were computed using the method described in Ref. 31. In contrast to Fourier spectra, in which a signal is transformed from the time domain to the

frequency domain, wavelet spectra provide the frequency content

of a signal on a time resolved basis. The Mexican Hat wavelet was

used for the analysis. The wavelet spectra were phase averaged to

show the variation of frequency content in the velocity as a function of time within the jet pulsing cycle.

3 Numerical Simulations

Numerical simulations were conducted utilizing version 6.3.26

of the finite-volume code FLUENTTM 32. The three dimensional

computational domain includes a single passage. A uniform velocity inflow condition is specified 1.9Cx upstream of the blade leading edge in the flow direction. The inlet flow angle is set to 33 deg

based on an inviscid calculation of the full cascade used in the

experiment 27. This angle agrees with the experimentally measured inlet angle to within the experiment uncertainty. The exit

boundary is located 3.8Cx downstream of the trailing edges in the

flow direction. In the spanwise direction, the domain includes one

VGJ. The boundary conditions on the sides of the passage are

periodic. The full length of the hole is included in the simulations,

allowing the jet velocity profile to develop before entering the

main domain 29. A uniform velocity boundary condition is

specified at the hole inlet during jet blowing. Computation was

continued until no variation cycle-to-cycle was reached.

Convergence was established when 1 residuals reduced to a

value 10−5, 2 no change was observed in any field results, and

3 the mass imbalance was less than 0.01%.

Large eddy simulations LESs were done using the dynamic

subgrid-scale kinetic energy model in FLUENT, which is based on

the model proposed by Kim and Menon 33. In this model, a

separate transport equation is solved for subgrid-scale kinetic energy. The model constants are determined dynamically. The details of the implementation of this model in FLUENT and its validation are given by Kim 34.

3.1 Code Validation. Three different grids were tested as

shown in Table 3. Grid 37 was selected. Details of the grid structure are in Ref. 35. To accurately represent structures in the near

wall region for LES, the recommended values are y+2, x+

50–150, and z+15–40 see Ref. 36. Grid 37 produced Cp

versus s/Ls results in closest agreement to the experimental data

and therefore was chosen for further computation. Computations

with grid 37 were then run for different time steps 0.0005,

0.0001, and 0.00005s and time step=0.0001 s was selected since

not much improvement was achieved using the smaller step

0.00005 s.

4 Results

4.1 Re= 25,000. Pressure profiles for cases with Re=25,000

and steady VGJ blowing are shown in Fig. 2. The inviscid profile

for the L1A airfoil is shown as a common reference for comparing

results between figures. At high Reynolds numbers e.g., Re

=300,000, experimental results agree with the inviscid line, as

shown in Ref. 28. The low peak followed by a plateau in the

case without jets in Fig. 2 indicates separation without reattachment. As in the low freestream turbulence cases of Ref. 29at

this Reynolds number, blowing with B=2.0 or lower has no effect

on separation. When B is increased to 2.5 or 3.0, there are some

signs of reattachment, but the Cp profile remains significantly

different from the inviscid profile.

Figures 3 shows results for cases with pulsed jets and F=0.14.

As in the low freestream turbulence cases, this frequency is too

low to be very effective. The jets have some effect on Cp when

B=3.0 and this effect increases at higher duty cycle. The total

pressure losses for these cases are shown in Fig. 3b. A high

Reynolds number Re=300,000 baseline case from Ref. 28is

shown for comparison. The loss, , is shown as a function of

distance across the cascade, , normalized on the blade spacing

L. The origin, =0, corresponds to the location downstream of

the trailing edge of the center blade in the cascade B4 in Fig.

1a in the design flow direction. The pulsed jets reduce the loss

significantly and shift the loss peaks to the right. This indicates

that the separation bubble must be thinner, particularly when B

=3.0, and that flow turning is increased. The loss peak at B=3.0 is

narrower for the D=50% case than for the lower duty cycle, in

agreement with the larger drop in Cp near the trailing edge for this

case. Even in this best case, the loss peaks are significantly higher

and to the left of the high Re case.

Results with F=0.28 are shown in Fig. 4. The jets begin to have

an effect when B=1.5, and significant separation control is apparent when B=2.0 or higher, particularly with D=50%. The jets are

not as effective when D=10%. The loss results show the VGJs at

Table 3 Grids 3, 35, and 37 used in this investigation

Grid nos. Size cells

Number

of grids in

z direction y+ z+ x+

3 1,500,000 15 0.5 12.6 1–100

35 5,900,000 30 0.5 6.3 0.4–52

37 11,900,000 54 0.5 0.4–3.5 0.4–52

0 0.2 0.4 0.6 0.8 1

0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

Steady,B=1.0

Steady,B=2.0

Steady,B=2.5

Steady,B=3.0

Inviscid

Fig. 2 Cp profiles for steady blowing, Re=25,000 cases

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NASA/CR—2012-217415 179
F=0.28 have a strong effect. The duty cycle effect is less apparent

than in the Cp results.

Figure 5 shows Cp results with F=0.56. The VGJs have some

effect even at low blowing ratios, and when B=1.0 the boundary

layer appears to reattach. The results do not appear to depend

significantly on the duty cycle. This is consistent with the low

freestream turbulence results, which showed that when F=0.56

the pulses occur frequently enough to control separation even at

low duty cycles. Figure 6 shows results with F=1.12. Results are

similar to those with F=0.56. Separation is controlled when B

=0.75, and increasing B further does not appear to have much

effect. Losses for the F=0.56 and 1.12 cases are shown in Fig. 7.

The losses are greatly reduced when separation is controlled, and

with a few exceptions the loss results are consistent with the Cp

results.

Integrated Cp and loss results are shown in Fig. 8. The Cp

difference between the pressure and suction sides of the passage is

integrated in the axial direction to produce a quantity proportional

to the lift on the airfoil,

Cpint = 

0

Cx Cpsuction − Cppressuredx

Cx

1

and this quantity is normalized using the same integrated quantity

for the inviscid profile. The integrated total pressure loss,

int = 

−L/2

L/2 d

L

+ m ̇ j

m ̇ 1

Pj − PTe

PT − PS 2

is an average of the loss profile result across one blade spacing

centered on the peak corresponding to blade B4. As explained in

0 0.2 0.4 0.6 0.8 1

0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

10%,B=2.0

10%,B=3.0

50%,B=1.0

50%,B=2.0

50%,B=3.0

Inviscid

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.5

1

1.5

φ/Lφ

ψ

B3

B4

B5 B6

No control

10%,B=2.0

10%,B=2.5

10%,B=3.0

50%,B=2.0

50%,B=3.0

Steady,B=3.0

High Re

Fig. 3 Pressure results for F=0.14, Re=25,000 cases: „a... Cp

and „b... total pressure loss

0 0.2 0.4 0.6 0.8 1

0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

10%,B=1.0

10%,B=1.5

10%,B=2.0

10%,B=3.0

50%,B=1.0

50%,B=2.0

50%,B=2.5

50%,B=3.0

Inviscid

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.5

1

1.5

φ/Lφ

ψ

B3

B4

B5 B6

No control

10%,B=1.5

10%,B=2.0

10%,B=3.0

50%,B=2.0

50%,B=3.0

High Re

Fig. 4 Pressure results for F=0.28, Re=25,000 cases: „a... Cp

and „b... total pressure loss

0 0.2 0.4 0.6 0.8 1

0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

10%,B=0.5

10%,B=0.75

10%,B=1.0

10%,B=1.5

10%,B=2.0

10%,B=2.5

10%,B=3.0

50%,B=0.5

50%,B=1.0

50%,B=2.0

Inviscid

Fig. 5 Cp profiles for F=0.56, Re=25,000 cases

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NASA/CR—2012-217415 180
Ref. 29, the loss associated with the jets themselves is included

in the integrated loss. With steady jets the lift increases slowly as

blowing ratio is increased but never achieves the inviscid case

value. The lift increases at each blowing ratio as F is increased

and is higher at the higher duty cycle. With F=0.56 and 1.12, the

lift increases rapidly with B, reaching the inviscid value when B

=1, and then remains constant as B is increased further. In agreement with the lift, the integrated loss decreases with B from a high

value of 0.55 without flow control to about 0.22, with the best

results at the higher pulsing frequencies. These values compare

with a high Re value of about 0.15. The losses with F=0.28 are

somewhat lower than in the corresponding low TI cases of Volino

et al. 29, but otherwise the high and low TI results are very

similar. The change in exit flow angle is inferred from the shift in

the location of the loss peaks from /L=0 in Figs. 3b, 4b,

and 7, and is shown as the difference from the high Re case angle

in Fig. 8c. The location of the peak in the high Re case corresponds to an exit flow angle of 54 deg, which agrees within the

experimental uncertainty with the design exit flow angle of 57 deg

corresponding to the design Zweifel=1.35. The design exit flow

angle is lower than the 60 deg design metal angle of Table 1.

Without flow control, turning is reduced by 16 deg compared with

the high Re case. In the best controlled case, about 11 deg of

turning is recovered, but there is still a 5.5 deg difference from the

high Re case due to the much thicker boundary layer at low Re

even with reattachment.

Time averaged velocity profiles for the baseline case and four

cases with pulsed jets are shown in Fig. 9. The top row shows the

mean velocity at the six streamwise stations noted in Table 2, and

the lower row shows the rms fluctuating streamwise velocity, u.

Without flow control, the boundary layer has separated by Station

1 and the separation bubble grows at the downstream stations. The

peak in u is in the shear layer far from the wall. With flow

control, it appears there may be a small separation bubble at the

0 0.2 0.4 0.6 0.8 1

0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

10%,B=0.5

10%,B=0.75

10%,B=1.0

10%,B=1.5

10%,B=2.0

Inviscid

Fig. 6 Cp profiles for F=1.12, Re=25,000 cases

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.5

1

1.5

φ/Lφ

ψ

B3

B4

B5 B6

No control

10%,B=0.75,F=0.56

10%,B=1.0,F=0.56

10%,B=2.0,F=0.56

50%,B=0.75,F=0.56

50%,B=1.0,F=0.56

50%,B=2.0,F=0.56

10%,B=1.0,F=1.12

10%,B=2.0,F=1.12

High Re

Fig. 7 Total pressure loss profiles for F=0.56 and F=1.12, Re

=25,000 cases

0 0.5 1 1.5 2 2.5 3

0.8

0.85

0.9

0.95

1

1.05

1.1

B

Cpint/Cpinv

Steady jets

F=0.14, D=10

F=0.14, D=50

F=0.28, D=10

F=0.28, D=50

F=0.56, D=10

F=0.56, D=50

F=1.12, D=10

0 0.5 1 1.5 2 2.5 3

0

0.1

0.2

0.3

0.4

0.5

0.6

B

ψint

High Re

No jets

Steady jets

F=0.14 D=10

F=0.14 D=50

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

F=1.12 D=10

0 0.5 1 1.5 2 2.5 3

0

2

4

6

8

10

12

14

16

B

∆θ [deg]

No jets

Steady jets

F=0.14 D=10

F=0.14 D=50

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

F=1.12 D=10

Fig. 8 Integrated pressure results for Re=25,000 cases: „a...

ratio of lift to lift in inviscid case, „b... total pressure loss, and „c...

change in exit flow angle from high Re case

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NASA/CR—2012-217415 181
third and fourth stations, in agreement with the short plateau in Cp

at this location in Figs. 4a and 5, but it is much smaller than in

the baseline case, and the boundary layer appears attached at the

downstream stations. With F=0.28, increasing the blowing ratio

from B=2 to 3 moves the u peak toward the wall, indicating a

thinner bubble. The same trend is observed as B in increased from

1 to 2 with F=0.56. Increasing F from 0.28 to 0.56 results in

significantly lower u in the outer part of the boundary layer,

which is explained below.

Figure 10 shows the phase averaged mean velocity profiles for

the cases with F=0.28, D=10%, and B=2.0, and with F=0.56,

D=10%, and B=2.0. The columns correspond to the six streamwise stations, and rows are for different phases in the pulsing

cycle. A small separation bubble appears to form at station 3 in the

F=0.28 case, but is suppressed for several phases starting at t/T

=0.417. The same is true at station 4, with the suppression of the

separation beginning at about t/T=0.5. At station 5, the boundary

layer appears to reattach at t/T=0.583 and remain attached until

t/T=0.917. At station 6, reattachment occurs at about t/T

=0.667, and the boundary layer remains attached until t/T

=0.167 in the next cycle.

In the F=0.56 case, there is less variation in the mean profile

during the cycle. Variation in the velocity as the boundary layer

separates and reattaches results in significant u, and as the variation is reduced at higher F, u decreases. This is particularly apparent at y /Ls0.1. As the separation bubble grows, fluid is

forced away from the wall and the velocity at y /Ls0.1 increases. When the boundary layer reattaches, the velocity at

y /Ls0.1 drops. The result is high u in this region for the F

=0.28 case in Fig. 9. The boundary layer appears to approach

separation in the F=0.56 case for much of the cycle at station 4,

particularly between t/T=0.5 and 0.708, and at t/T=0.75 to 0.917

at station 5 and t/T=0.083 at station 6. At other times, separation

appears controlled.

Wavelet spectra help to explain the velocity profile results. Figure 11 shows wavelet spectra for the F=0.28, D=10%, and B

=2.0 case. The six plots in the figure correspond to the six streamwise stations. In each plot, the horizontal axis shows dimensionless frequency, fLs/Ue on a log scale, and the vertical axis shows

dimensionless time, t/T, for one pulsing cycle. Power spectral

density is computed from instantaneous velocity data at all y locations and is shown in Fig. 11 for the y locations corresponding

to maximum time averaged u at each station. The contours show

the power spectral density premultiplied by frequency and normalized by Ue

2

. The color scale is the same for all plots. The VGJ

creates a disturbance at the beginning of a pulse, and the leading

edge of this disturbance is visible as high contours centered at

t/T=0.16, 0.25, 0.31, 0.38, 0.46, and 0.54 at stations 1–6, respectively. The arrival times at each station indicate that the leading

edge of the disturbance convects along the surface at about 0.6

times the local freestream velocity. A second peak appears centered at t/T=0.24, 0.42, 0.54, 0.71, 0.93, and 0.06 at stations 1–6,

respectively. These peaks are believed to result when the trailing

edge of the VGJ disturbance passes, and the times indicate a convection speed of about 0.3 times the freestream velocity. Comparison of Figs. 10 and 11 shows that at each station, the separation

bubble is thickest at about the time the disturbance first arrives.

Shortly after this, the velocity profile appears reattached. The profile remains attached at each station until about the time when the

second peak appears in Fig. 11, indicating that the disturbance has

past. The separation bubble then begins to grow almost immediately. Since the leading and trailing edges of the disturbance move

at different velocities, the duration of the disturbance increases at

the downstream stations. Hence, the duration of the separation

bubble decreases at the downstream stations. If the pulsing fre0 1 0 1 0 1 0 1 0 1 0 1 0

0.1

0.2

U/Ue

y L/ s

St 12 3 4 5 6

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0

0.1

0.2

u′/Ue

y/Ls

No jets F=0.28 B=2.0 F=0.28 B=3.0 F=0.56 B=1.0 F=0.56 B=2.0

Fig. 9 Time averaged velocity profiles at six streamwise stations for Re=25,000 cases with no jets and jets with D=10%:

„top... mean velocity and „bottom... rms velocity

0

0.1

0.083

t/T St. 12 3 4 5 6

0.083

0

0.1

0.167

0

0.1

0.25

0

0.1

0.333

0

0.1

0.417

0

0.1

0.5 y/Ls

0

0.1

0.583

0

0.1

0.667

0

0.1

0.75

0

0.1

0.833

0

0.1

0.917

0 1 0 1 0 1 0 1 0 1 0 1 0

0.1

1

U/Ue

Fig. 10 Phase averaged mean velocity profiles for Re=25,000

cases, columns for six streamwise stations, rows for phases in

pulsing cycle: black „blue...—F=0.28, D=10%, B=2.0; gray

„red...—F=0.56, D=10%, B=2.0

t/T

St. 1

0.1 1 10 0

0.2

0.4

0.6

0.8

1

2

0.1 1 10

3

f⋅Ls

/Ue

0.1 1 10

4

0.1 1 10

5

0.1 1 10

6

0.1 1 10

Fig. 11 Wavelet spectra computed at y locations of maximum

u in time averaged profiles and shown as function of time and

frequency at six streamwise stations, F=0.28, D=10%, B=2.0,

Re=25,000, solid white line is leading edge of disturbance,

dashed white line is trailing edge

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NASA/CR—2012-217415 182
quency were increased, disturbances would follow each other

more closely in time, and if the frequency were sufficiently high,

the leading edge of each disturbance would overtake the trailing

edge of the preceding disturbance, thereby eliminating the period

of separated flow. Based on the convection speeds observed in

Fig. 11, F=0.5 would be sufficient to eliminate separation at the

trailing edge.

Low frequency peaks appear in Fig. 11 at the pulsing frequency,

fLs/Ue=0.7, at all stations. A higher frequency peak at fLs/Ue

=10 is also visible, particularly at stations 1 and 2. The higher

frequency is likely related to shear layer transition and matches

the frequency peak observed in transition without flow control in

Ref. 28. Figure 12 shows the contours of the dimensionless,

premultiplied wavelet spectra as a function of distance from the

wall on the vertical axis and dimensionless frequency on the horizontal axis. The columns correspond to stations 2–6, and the rows

to different times during the cycle. The peak at fLs/Ue=0.7 appears at all times. Higher frequencies extending beyond fLs/Ue

=10 are also present. The high frequency fluctuations have their

highest magnitude in the shear layer when the boundary layer is

most separated and are suppressed when the VGJs cause reattachment. Another peak at fLs/Ue=2 and y /Ls between 0.11 and 0.16

appears in the shear layer at t/T=0.417, 0.5, and 0.583 at stations

4–6, respectively. These are the arrival times of the VGJ disturbance at these stations.

Increasing the blowing ratio from 2 to 3 with F=0.28 increases

the magnitude of the disturbances, but otherwise does not change

the behavior observed in Figs. 9–12. The effect of pulsing frequency is shown in Fig. 13 for the F=0.56, D=10%, and B=2.0

case in the same format as Fig. 11. A peak frequency of about

fLs/Ue=10 is still present at stations 1 and 2, as in Fig. 11. High

magnitudes are also visible at all stations at the pulsing frequency,

which is now doubled to about fLs/Ue=1.4. In the format of Fig.

12 for this case not shown, the arrival of the VGJ disturbance is

again marked by a spectral peak in the shear layer at a frequency

of about fLs/Ue=2. In Fig. 13, the center of the spectral peaks

which presumably correspond to the first arrival of the VGJ disturbance appear at stations 1–6 at t/T=0.17, 0.34, 0.5, 0.67, 0.94,

and 0.1 respectively. These dimensionless times are greater than

those in Fig. 11 since the convection velocities are about the same

in both cases but the pulsing period is halved. Comparing Figs. 13

and 10, the separation bubble begins to grow at each station after

the passing of the trailing edge of the disturbance, and the bubble

is suppressed about 0.1T after the arrival of the next disturbance.

This is the same behavior observed at F=0.28, but there is less

time available for the bubble to grow at the higher frequency, so

there is less variation of the velocity profile shape with time at

F=0.56. The reduced separation at higher frequency is reflected in

the lower losses and the Cp profiles, which are closer to the inviscid solution.

4.2 Re= 50,000. Pressure profiles for the cases with steady

VGJs and Re=50,000 are shown in Fig. 14. The jets become

effective when B1.5. Results for pulsed jets with F=0.14 are

shown in Fig. 15. In terms of Cp, the jets begin to have an effect

when B=0.75 but do not appear to fully control separation until

B=1.5. Between B=0.75 and 1.5, a higher duty cycle appears to

help, but at B=1.5 and above, the results are about the same with

D=10% or 50%. The total pressure loss profiles agree with the Cp

results. In the best controlled case, the loss peaks are still significantly higher and to the left of the high Re comparison case, but

the difference is not as great as in the Re=25,000 cases above.

Figure 16 shows results with F=0.28. The jets have some effect

with B=0.5 and 0.75, and the effect increases at higher duty cycle.

With B=1.0 and higher, separation is controlled at both duty

cycles. The loss peak is slightly narrower with D=50%. The Cp

profiles show a small separation bubble still appears to remain at

s/Ls=0.7. With F=0.56, as shown in Fig. 17, blowing with B

=0.5 has some effect, particularly with D=50%, and with B

=0.75 and D=10% the separation is controlled. Increasing B or D

further has no additional effect. As in the lower frequency cases, a

small separation bubble remains at s/Ls=0.7. The integrated Cp

St 2

0

0.2

0

0.2

0

0.2

0

0.2

0

0.2

y/Ls

0

0.2

0

0.2

0

0.2

0

0.2

0

0.2

0

0.2

0.1 1 10 0

0.2

3

0.1 1 10

4

f⋅Ls/Ue

0.1 1 10

5

0.1 1 10

6

0.083

t/T

0.167

0.25

0.333

0.417

0.5

0.583

0.667

0.75

0.833

0.917

1

0.1 1 10

Fig. 12 Wavelet spectra as function of distance from wall and

frequency at stations 2–6 „columns... and various phases

„rows..., F=0.28, D=10%, B=2.0, Re=25,000

t/T

St. 1

0.1 1 10 0

0.2

0.4

0.6

0.8

1

2

0.1 1 10

3

f⋅Ls

/Ue

0.1 1 10

4

0.1 1 10

5

0.1 1 10

6

0.1 1 10

Fig. 13 Wavelet spectra computed at y locations of maximum

u in time averaged profiles and shown as function of time and

frequency at six streamwise stations, F=0.56, D=10%, B=2.0,

Re=25,000, solid white line is leading edge of disturbance,

dashed white line is trailing edge

0 0.2 0.4 0.6 0.8 1

0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

Steady,B=0.5

Steady,B=0.75

Steady,B=1.0

Steady,B=1.5

Steady,B=2.0

Inviscid

Fig. 14 Cp profiles for steady blowing, Re=50,000 cases

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NASA/CR—2012-217415 183
and loss results are shown in Fig. 18. The trends are the same as

in the Re=25,000 cases of Fig. 8 and the low TI cases of Volino et

al. 29, but effective separation control and loss reduction is possible with somewhat lower B and F as Re or TI is increased. The

reduction in flow turning compared with the high Re case is

shown in Fig. 18c. Separation control increases the flow turning

by about 12 deg, reducing the difference from the high Re case to

2 deg.

Time averaged velocity profiles for some of the Re=50,000

cases are shown in Fig. 19. A large separation bubble without

reattachment is present without flow control. Pulsed VGJs largely

eliminate the bubble, although a small separation is still present at

station 3, in agreement with the small plateau in the Cp profiles of

Figs. 16 and 17 at s/Ls=0.7. Figure 20 shows phase averaged

velocity profiles. With F=0.28, there is a small separation at

t/T=0.417 at station 4, at t/T=0.5 at station 5, and at t/T

=0.583 at station 6. Figure 21 shows the wavelet spectra for this

case in the format of Fig. 11. The frequencies and times of the

peaks in Figs. 21 and 11 are similar since the pulsing frequency,

F=0.28, is the same for both cases. As in the lower Re case, the

boundary layer reattaches shortly after the appearance of the VGJ

disturbance at each station. In contrast to the lower Re case, the

boundary layer remains attached for about 0.2T after the disturbance has passed, resulting in attached flow for most of the pulsing cycle. The F=0.56, D=10%, B=0.75 case in Fig. 20 shows

little variation of the velocity profiles during the pulsing cycle.

The spectra for this case, shown in Fig. 22, show nearly a continuous disturbance through the cycle. There may be some signs

of separation with a thin bubble at station 3, but at the downstream stations the near overlap of the disturbance period between

consecutive pulses combined with the lesser tendency toward

separation at Re=50,000 than at 25,000, results in an attached

boundary layer for the full cycle.

4.3 Computational Results. Figure 23 shows Cp for the Re

=50,000, B=1, D=10%, F=0.28 case. The spike at the suction

peak is at the jet location. In agreement with the experimental

data, the CFD shows a small separation bubble occurs but the flow

remains reattached over most of the airfoil. Figure 24 shows time

averaged mean velocity at six streamwise stations comparing the

CFD results with the experimental data. The agreement is good

upstream, but differences are apparent at the downstream stations,

particularly near the wall. The CFD predicts an exit flow angle of

57 deg both in the present case and in a high Re case. The exit

angle difference between the high Re and controlled Re=50,000

case was also small in the experimental result in Fig. 18c. 57 deg

agrees with the design exit flow angle but is a few degrees higher

than the experimental value. Figure 25 shows the Q-criterion,

which is used to illustrate vortices 32, colored by axial velocity,

Vx, at different times in the cycle. At the beginning of the blowing,

there is no separation bubble present near the trailing edge of the

airfoil. This is in contrast with the corresponding low TI case see

Ref. 35. However, there is a small bubble located midway between the jet and the trailing edge. During blowing that separated

region travels downstream while the flow is attached along most

0 0.2 0.4 0.6 0.8 1

0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

10%,B=0.25

10%,B=0.5

10%,B=0.75

10%,B=1.0

10%,B=1.5

10%,B=2.0

50%,B=0.25

50%,B=0.5

50%,B=0.75

50%,B=1.0

50%,B=1.5

50%,B=2.0

Inviscid

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.5

1

1.5

φ/Lφ

ψ

B3

B4

B5 B6

No control

10%,B=1.0

10%,B=1.5

10%,B=2.0

50%,B=1.0

50%,B=1.5

High Re

Fig. 15 Pressure results for F=0.14, Re=50,000 cases: „a... Cp

and „b... total pressure loss

0 0.2 0.4 0.6 0.8 1

0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

10%,B=0.25

10%,B=0.5

10%,B=0.75

10%,B=1.0

10%,B=1.5

50%,B=0.25

50%,B=0.5

50%,B=0.75

50%,B=1.0

Inviscid

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.5

1

1.5

φ/Lφ

ψ

B3

B4

B5 B6

No control

10%,B=0.75

10%,B=1.0

50%,B=0.5

50%,B=0.75

50%,B=1.0

High Re

Fig. 16 Pressure results for F=0.28, Re=50,000 cases: „a... Cp

and „b... total pressure loss

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NASA/CR—2012-217415 184
of the airfoil. Right after shutdown of the jet, the separation region

is smaller. Also at that time the shear layer is highly energized.

Near the end of the cycle the flow is attached at the trailing edge,

but a small separated region begins to appear again upstream.

These results are consistent with the experimental velocity profiles

of Fig. 20, which show a small separation bubble moving down

the airfoil between wakes, but attached flow at all locations for

most of the cycle.

5 Discussion

The VGJ disturbances convect downstream in the shear layer

over the separation bubble, and if they are of sufficient amplitude,

cause the boundary layer to attach at a given streamwise location

shortly after the arrival of the disturbance. The boundary layer

tends to remain attached until after the passage of the trailing edge

of the disturbance, although at the lowest Reynolds numbers there

may be some tendency, as shown in Fig. 10, to separate earlier.

After the disturbance passes, the separation bubble begins to regrow. This happens almost immediately at very low Reynolds

number and more slowly at higher Re. The trailing edge of each

disturbance travels slower than the leading edge, resulting in a

stretching of the disturbance in the streamwise direction as it

moves downstream and a longer duration of the disturbance at

downstream locations. The result is a shorter period between disturbances at downstream stations and less separated flow. Increasing the pulsing frequency reduces the period between disturbances

at all locations and results in better separation control.

The present results, in agreement with all the previous VGJ

studies referenced above, show that pulsed jets are more effective

than steady jets. The disturbance created when a jet is first turned

on is more effective than any steady blowing that may follow.

Velocity results from cases with low freestream turbulence 37

show that when pulsing with D=50%, the boundary layer may

reseparate during the period of steady blowing. Increasing the

duty cycle from 10% to 50% can, however, provide better separation control, particularly at low pulsing frequencies, as shown

above and in Ref. 29. Velocity profiles and spectra from the low

0 0.2 0.4 0.6 0.8 1

0

0.5

1

1.5

2

2.5

s/Ls

Cp

No jets

10%,B=0.25

10%,B=0.5

10%,B=0.75

10%,B=1.0

50%,B=0.25

50%,B=0.5

50%,B=0.75

Inviscid

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.5

1

1.5

φ/Lφ

ψ

B3

B4

B5 B6

No control

10%,B=0.5

10%,B=0.75

10%,B=1.0

50%,B=0.5

50%,B=0.75

High Re

Fig. 17 Pressure results for F=0.56, Re=50,000 cases: „a... Cp

and „b... total pressure loss

0 0.5 1 1.5 2 0.8

0.85

0.9

0.95

1

1.05

B

Cpint/Cpinv

Steady jets

F=0.14, D=10

F=0.14, D=50

F=0.28, D=10

F=0.28, D=50

F=0.56, D=10

F=0.56, D=50

0 0.5 1 1.5 2 0

0.1

0.2

0.3

0.4

0.5

0.6

B

ψint

High Re

No jets

F=0.14 D=10

F=0.14 D=50

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

0 0.5 1 1.5 2 0

2

4

6

8

10

12

14

16

B

∆θ [deg]

No jets

F=0.14 D=10

F=0.14 D=50

F=0.28 D=10

F=0.28 D=50

F=0.56 D=10

F=0.56 D=50

Fig. 18 Integrated pressure results for Re=50,000 cases: „a...

ratio of lift to lift in inviscid case, „b... total pressure loss, and „c...

change in exit flow angle from high Re case

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NASA/CR—2012-217415 185
TI, D=50% cases show a second effective disturbance occurs

each cycle when the VGJs turns off, thereby doubling the disturbance frequency above the pulsing frequency. In the D=10%

cases, the disturbances created when the jets turn on and off occur

so close together in time that they effectively constitute a single

disturbance.

The present results are very similar to the low TI results of

Volino et al. 29. High TI does help to control separation, allowing VGJ flow control with slightly lower amplitude and frequency

jets in some cases, but the physics of the separation and flow

control appears to be the same at high and low TI. Increasing the

VGJ velocity and frequency help to reduce the separation bubble

extent both temporally and spatially, and this increases lift and

reduces total pressure losses. Once the separation is largely eliminated, increasing the blowing ratio or pulsing frequency further

provides no additional benefit.

6 Conclusions

The effect of vortex generator jets on the flow over the very

high lift L1A airfoil was studied under high freestream turbulence

conditions. Reynolds numbers based on suction surface length and

nominal exit velocity of 25,000 and 50,000 were considered.

Without flow control, the boundary layer separated and did not

reattach. Flow control with VGJs was possible even at Re

=25,000. In agreement with previous studies, pulsed jets were

more effective than steady jets. Pulsing with a dimensionless frequency of F=0.56 or higher allowed for separation control with a

blowing ratio of at most 1.0 and 10% duty cycle. Effective separation control resulted in a 20% increase in lift and up to a 60%

reduction in total pressure loss compared with baseline cases at

the same Reynolds number. Loss values still remain higher than in

high Reynolds number cases. Pulsed jets at lower frequencies pro0 1 0 1 0 1 0 1 0 1 0 1 0

0.1

0.2

U/Ue

y/Ls

St 12 3 4 5 6

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0

0.1

0.2

u′/Ue

y/Ls

No jets F=0.28 B=1.0 F=0.56 B=0.75

Fig. 19 Time averaged velocity profiles at six streamwise stations for Re=50,000 cases with no jets and jets with D=10%:

„top... mean velocity and „bottom... rms velocity

0

0.1

0.083

t/T St. 12 3 4 5 6

0.083

0

0.1

0.167

0

0.1

0.25

0

0.1

0.333

0

0.1

0.417

0

0.1

0.5 y L/ s

0

0.1

0.583

0

0.1

0.667

0

0.1

0.75

0

0.1

0.833

0

0.1

0.917

0 1 0 1 0 1 0 1 0 1 0 1 0

0.1

1

U/Ue

Fig. 20 Phase averaged mean velocity profiles for Re=50,000

cases, columns for six streamwise stations, rows for phases in

pulsing cycle: black „blue...—F=0.28, D=10%, B=1.0; gray

„red...—F=0.56, D=10%, B=0.75 t/T

St. 1

0.1 1 10 0

0.2

0.4

0.6

0.8

1

2

0.1 1 10

3

f⋅Ls

/Ue

0.1 1 10

4

0.1 1 10

5

0.1 1 10

6

0.1 1 10

Fig. 21 Wavelet spectra computed at y locations of maximum

u in time averaged profiles and shown as function of time and

frequency at six streamwise stations, F=0.28, D=10%, B=1.0,

Re=50,000; the solid white line is the leading edge of disturbance, and the dashed white line is the trailing edge

t/T

St. 1

0.1 1 10 0

0.2

0.4

0.6

0.8

1

2

0.1 1 10

3

f⋅Ls

/Ue

0.1 1 10

4

0.1 1 10

5

0.1 1 10

6

0.1 1 10

Fig. 22 Wavelet spectra computed at y locations of maximum

u in time averaged profiles and shown as function of time and

frequency at six streamwise stations, F=0.56, D=10%, B=0.75,

Re=50,000; the solid white line is leading edge of disturbance,

and the dashed line is trailing edge

Fig. 23 Cp profile for B=1, F=0.28, D=10%, Re=50,000

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NASA/CR—2012-217415 186
vided partial separation control and some loss reduction and could

be improved somewhat by increasing the blowing ratio or duty

cycle.

Phase averaged velocity profiles and wavelet spectra show the

boundary layer intermittently reattaching as disturbances pass and

then separating between disturbances. Increasing the pulsing frequency reduces the time available for separation. When the time

available is sufficiently small, the boundary layer remains attached at all times. At Re=25,000, separation was nearly fully

controlled for the full pulsing cycle when F=0.5. Higher frequency pulsing provided little additional benefit. At higher Reynolds numbers, the separation bubble grows more slowly, so

lower pulsing frequencies can be tolerated.

Computational results from a large eddy simulation show reattachment in agreement with the experiments for a case with effective VGJ flow control. The LES results allow visualization of

structures such as vortices, providing insight into the flow behavior.

High freestream turbulence makes the boundary layer less

likely to separate and allows for separation control with slightly

lower blowing ratios and pulsing frequencies than in low TI cases.

The differences observed between high and low TI were small,

however, and the physics of the separation and flow control appear to be the same in both cases.

Acknowledgment

This work was sponsored by the National Aeronautics and

Space Administration under Grant No. NNC07IA10I. The grant

monitors were Dr. Anthony Strazisar and Dr. James Heidmann of

the NASA Glenn Research Center. The support of the United

States Naval Academy Technical Support Department Shop and

Fluids Laboratory is greatly appreciated. We greatly appreciate the

computer time provided for us by the Ohio Super Computer

OSC. The OSC Computer Cluster has been made available as

part of the Center’s mission to support Ohio Universities.

Nomenclature

B  blowing ratio, maximum jet velocity/local

freestream velocity

Cp  2PT−P/Ue

2

, pressure coefficient

Cpint  pressure coefficient integrated in axial direction

Cpinv  Cpint for inviscid flow through cascade

Cx  axial chord length

D  duty cycle, fraction of time valve is open

F  fLj−te /Uave, dimensionless frequency

f  frequency

Lj−te  distance from VGJs to trailing edge

Ls  suction surface length

Fig. 24 U/Ue for B=1, F=0.28, D=10%, Re=50,000

Fig. 25 Q-criterion contours colored by axial velocity, B=1, F

=0.28, D=10%, Re=50,000, at different times in the cycle

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NASA/CR—2012-217415 187
L  blade spacing pitch

m ̇ j  jet mass flow rate

m ̇ 1  main flow mass flow rate

P  pressure

Pj  jet plenum pressure

PS  upstream static pressure

PT  upstream stagnation pressure

PTe  downstream stagnation pressure

Re  UeLs/	, exit Reynolds number

s  streamwise coordinate, distance from leading

edge

T  period of jet pulsing cycle

t  time

Uave  average freestream velocity between VGJs and

trailing edge

Ue  nominal exit freestream velocity, based on inviscid solution

u  rms fluctuating streamwise velocity

x  axial distance from leading edge

x+  grid spacing in axial direction in viscous units

y+  location of closest grid point to wall in viscous

units

z+  grid spacing in spanwise direction in viscous

units

  coordinate along blade spacing, normal to axial

chord

	  kinematic viscosity

  density

  PT−PTe/PT−PS, total pressure loss

coefficient

int  total pressure loss integrated over blade

spacing

References

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AGARD Lecture Series 167.

2Mayle, R. E., 1991, “The Role of Laminar-Turbulent Transition in Gas Turbine

Engines,” ASME J. Turbomach., 113, pp. 509–537.

3Sharma, O.P., Ni, R.H., and Tanrikut, S., 1994, “Unsteady Flow in Turbines,”

AGARD Lecture Series 195, Paper No. 5.

4Bons, J. P., Sondergaard, R., and Rivir, R. B., 2001, “Turbine Separation

Control Using Pulsed Vortex Generator Jets,” ASME J. Turbomach., 123, pp.

198–206.

5Volino, R. J., and Hultgren, L. S., 2001, “Measurements in Separated and

Transitional Boundary Layers Under Low-Pressure Turbine Airfoil Conditions,” ASME J. Turbomach., 123, pp. 189–197.

6Volino, R. J., 2010, “Separated Flow Measurements on a Highly Loaded LowPressure Turbine Airfoil,” ASME J. Turbomach., 132, p. 011007.

7Volino, R. J., 2002, “Separated Flow Transition Under Simulated LowPressure Turbine Airfoil Conditions: Part 1—Mean Flow and Turbulence Statistics,” ASME J. Turbomach., 124, pp. 645–655.

8Volino, R. J., 2002, “Separated Flow Transition Under Simulated LowPressure Turbine Airfoil Conditions: Part 2—Turbulence Spectra,” ASME J.

Turbomach., 124, pp. 656–664.

9Praisner, T. J., and Clark, J. P., 2007, “Predicting Transition in

Turbomachinery—Part 1: A Review and New Model Development,” ASME J.

Turbomach., 129, pp. 1–13.

10Bons, J.P., Hansen, L.C., Clark, J.P., Koch, P.J., and Sondergaard, R., 2005,

“Designing Low-Pressure Turbine Blades With Integrated Flow Control,”

ASME Paper No. GT2005-68962.

11Zhang, X. F., Vera, M., Hodson, H., and Harvey, N., 2006, “Separation and

Transition Control on an Aft-Loaded Ultra-High-Lift LP Turbine Blade at Low

Reynolds Numbers: Low-Speed Investigation,” ASME J. Turbomach., 128,

pp. 517–527.

12Bohl, D. G., and Volino, R. J., 2006, “Experiments With Three-Dimensional

Passive Flow Control Devices on Low-Pressure Turbine Airfoils,” ASME J.

Turbomach., 128, pp. 251–260.

13Volino, R. J., 2003, “Passive Flow Control on Low-Pressure Turbine Airfoils,”

ASME J. Turbomach., 125, pp. 754–764.

14Volino, R. J., 2003, “Separation Control on Low-Pressure Turbine Airfoils

Using Synthetic Vortex Generator Jets,” ASME J. Turbomach., 125, pp. 765–

777.

15Huang, J., Corke, T., and Thomas, F., 2003, “Plasma Actuators for Separation

Control on Low Pressure Turbine Blades,” AIAA Paper No. 2003-1027.

16Johnston, J. P., and Nishi, M., 1990, “Vortex Generator Jets—Means for Flow

Separation Control,” AIAA J., 28, pp. 989–994.

17Bons, J. P., Sondergaard, R., and Rivir, R. B., 2002, “The Fluid Dynamics of

LPT Blade Separation Control Using Pulsed Jets,” ASME J. Turbomach., 124,

pp. 77–85.

18Volino, R. J., and Bohl, D. G., 2005, “Structure of Oscillating Vortex Generator Jets,” Proceedings of the Fourth International Symposium on Turbulence

and Shear Flow Phenomena, Vol. 2, pp. 589–594.

19McQuilling, M., and Jacob, J., 2004, “Effect of Chord Location on Separation

Control With Vortex Generator Jets on Low Pressure Turbine Blades,” AIAA

Paper No. 2004-2205.

20Eldredge, R. G., and Bons, J. P., 2004, “Active Control of a Separating Boundary Layer With Steady Vortex Generating Jets—Detailed Flow Measurements,” AIAA Paper No. 2004-751.

21Gostelow, J. P., Walker, G. J., Solomon, W. J., Hong, G., and Melwani, N.,

1997, “Investigation of the Calmed Region Behind a Turbulent Spot,” ASME

J. Turbomach., 119, pp. 802–809.

22Schulte, V., and Hodson, H. P., 1998, “Prediction of the Becalmed Region for

LP Turbine Profile Design,” ASME J. Turbomach., 120, pp. 839–846.

23Bons, J. P., Reimann, D., and Bloxham, M., 2008, “Separated Flow Transition

on an LP Turbine Blade With Pulsed Flow Control,” ASME J. Turbomach.,

130, p. 021014.

24Clark, J. P., 2007, Private Communication, Air Force Research Laboratory.

25Zhang, X. F., and Hodson, H., 2005, “Combined Effects of Surface Trips and

Unsteady Wakes on the Boundary Layer Development of an Ultra-High-Lift

LP Turbine Blade,” ASME J. Turbomach., 127, pp. 479–488.

26Bons, J. P., Pluim, J., Gompertz, K., Bloxham, M., and Clark, J. P., 2008, “The

Application of Flow Control to an Aft-Loaded Low Pressure Turbine Cascade

With Unsteady Wakes,” ASME Paper No. GT2008-50864.

27Ibrahim, M., Kartuzova, O., and Volino, R. J., 2008, “Experimental and Computational Investigations of Separation and Transition on a Highly Loaded

Low Pressure Turbine Airfoil: Part 1—Low Freestream Turbulence Intensity,”

ASME Paper No. IMECE2008-68879.

28Volino, R. J., Kartuzova, O., and Ibrahim, M., 2008, “Experimental and Computational Investigations of Separation and Transition on a Highly Loaded

Low Pressure Turbine Airfoil: Part 2—High Freestream Turbulence Intensity,”

ASME Paper No. IMECE2008-68776.

29Volino, R. J., Kartuzova, O., and Ibrahim, M. B., 2009, “Experimental and

Computational Investigations of Low-Pressure Turbine Separation Control Using Vortex Generator Jets,” ASME Paper No. GT2009-59983.

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a Solid Boundary,” J. Fluid Mech., 12, pp. 388–396.

31Volino, R. J., 2005, “An Investigation of the Scales in Transitional Boundary

Layers Under High Freestream Turbulence Conditions,” Exp. Fluids, 38, pp.

516–533.

32Fluent, Inc., 2005, FLUENT 6.3 User Guide.

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Control Simulations for Highly Loaded Low Pressure Turbine Airfoil L1A

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Journal of Turbomachinery OCTOBER 2011, Vol. 133 / 041021-13

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NASA/CR—2012-217415 188
Ralph J. Volino

Department of Mechanical Engineering,

United States Naval Academy,

Annapolis, MD 21402-5042

e-mail: volino@usna.edu

Effect of Unsteady Wakes on

Boundary Layer Separation on a

Very High Lift Low Pressure

Turbine Airfoil

Boundary layer separation has been studied on a very high lift, low pressure turbine

airfoil in the presence of unsteady wakes. Experiments were done under low (0.6%) and

high (4%) freestream turbulence conditions on a linear cascade in a low speed wind

tunnel. Wakes were produced from moving rods upstream of the cascade. Flow coefficients were varied from 0.35 to 1.4 and wake spacing was varied from one to two blade

spacings, resulting in dimensionless wake passing frequencies F fLj-te /Uave (f is the

frequency, Lj-te is the length of the adverse pressure gradient region on the suction

surface of the airfoils, and Uave is the average freestream velocity) ranging from 0.14 to

0.56. Pressure surveys on the airfoil surface and downstream total pressure loss surveys

were documented. Instantaneous velocity profile measurements were acquired in the suction surface boundary layer and downstream of the cascade. Cases were considered at

Reynolds numbers (based on the suction surface length and the nominal exit velocity from

the cascade) of 25,000 and 50,000. In cases without wakes, the boundary layer separated

and did not reattach. With wakes, separation was largely suppressed, particularly if the

wake passing frequency was sufficiently high. At lower frequencies the boundary layer

separated between wakes. Background freestream turbulence had some effect on separation, but its role was secondary to the wake effect. DOI: 10.1115/1.4003232

1 Introduction

Partial loss of lift and higher aerodynamic losses can be caused

by boundary layer separation on low pressure turbine LPT airfoils e.g., Hourmouziadis 1, Mayle 2, and Sharma et al. 3.

In aircraft engines the lower density and therefore lower Reynolds

numbers at altitude can lead to a component efficiency drop of 2%

between takeoff and cruise in large commercial transport engines

and possibly as much as 7% in smaller engines operating at higher

altitudes 4,5. Separation can occur when airfoil loading is too

high because of the strong adverse pressure gradients on the suction surface. High loading is, however, desirable since it can be

used to reduce airfoil count, weight, and cost. The challenge is to

design high lift airfoils, which do not have separation problems.

This requires accurate prediction of suction side separation under

relevant conditions.

Separation can be strongly affected by wakes shed from the

airfoils in upstream stages in an engine. The velocity deficit and

elevated turbulence in periodic wakes help to suppress separation

and can cause a separated boundary layer to reattach. Hodson and

Howell 6described the mechanisms by which wakes promote

reattachment, including the “negative jet,” which results when the

velocity deficit in the wake causes the flow outside the wake to

accelerate and impinge on the surface, and the unsteadiness,

which promotes transition in the boundary layer. Following the

wake itself is a calmed period Gostelow et al. 7and Schulte and

Hodson 8 in which the boundary layer has low turbulence but is

resistant to separation. Numerous studies have considered the

wake effect in the LPT, including those listed in Hodson and

Howell 6, and more recent references in Bons et al. 9and

Pluim et al. 10. Examples include Schobeiri et al. 11, Öztürk

and Schobeiri 12, Jiang and Simon 13, and Mahallati and

Sjolander 14who all used the Pack B airfoil. Zhang and Hodson

15and Funazaki et al. 16used more highly loaded airfoils.

Many additional studies are available from these research groups

and others.

Airfoils can be designed with high resistance to separation, as

described by Praisner and Clark 17, and knowledge of wake

effects allows for designs with higher loading than would be possible under steady inflow conditions. Even with wakes, however, a

loading limit will always exist, above which separation will still

occur.

In the present study, a very highly loaded airfoil that exhibits

separation even in the presence of wakes is used. The airfoil,

known as the L1A, was designed at the Air Force Research Laboratory AFRL and is available on a limited basis from Clark 18.

Dimensions of the L1A as used in the present study are given in

Table 1. Based on the design calculations of Clark 18, the L1A

has a Zweifel coefficient of 1.35, which corresponds to 10%

higher loading than the “ultrahigh lift” airfoils described by Zhang

and Hodson 19and 17% higher loading than the Pack B airfoil.

The L1A is aft loaded, which is advantageous for reducing secondary flow losses at the endwalls, but makes the boundary layer

more prone to separation than a forward loaded blade, as documented in Bons et al. 9, Volino 20, Ibrahim et al. 21, and

Volino et al. 22. In cases without wakes and low Reynolds numbers, the boundary layer separates and does not reattach, in spite

of transition to turbulence in the shear layer over the separation

bubble. This result contrasts with the results of studies on less

aggressive airfoils e.g., Volino 23, which all showed reattachment after transition. The separation bubble on the L1A is about

four times thicker than that on the Pack B. The larger distance

from the shear layer to the wall on the L1A apparently prevents

the turbulent mixing in the shear layer from reaching the wall and

causing reattachment. The failure of the boundary layer to reattach

Contributed by the International Gas Turbine Institute IGTI of ASME for publication in the JOURNAL OF TURBOMACHINERY. Manuscript received September 3, 2010;

final manuscript received October 23, 2010; published online May 26, 2011. Editor:

David Wisler.

This material is declared a work of the US government and is not subject to

copyright protection in the United States. Approved for public release; distribution is

unlimited.

Journal of Turbomachinery JANUARY 2012, Vol. 134 / 011011-1

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NASA/CR—2012-217415 189
results in a 20% loss in lift and increases profile losses by up to a

factor of 7. At higher Reynolds numbers the separation bubble

closes, and for Re200,000 the separation bubble is small and

the boundary layer is attached over most of the airfoil.

The effect of wakes on the L1A boundary layer was studied by

Bons et al. 9, who considered a case with Re= 50,000 based on

the suction surface length and the nominal exit velocity from the

cascade, background freestream turbulence TI= 3%, and periodic

wakes produced with moving rods upstream of the airfoils. The

dimensionless frequency of the wake passing was F= fLj-te /Uave

= 0.34, where Lj-te is the length of the adverse pressure gradient

region on the suction surface and Uave is the average freestream

velocity over this distance. The length Lj-te is also the distance

from a row of vortex generator jet VGJ holes to the trailing

edge. Pulsed vortex generator jets were used by Bons et al. 9and

Volino et al. 24to control separation on the L1A. Although these

jets are not considered in the present work, the same dimensionless frequency, F, is used to describe the wake passing in the

present study and the jet pulsing in Refs. 9,24. Bons et al. 9

found that wakes reduced the separation bubble thickness significantly and reduced total pressure losses by more than 50%, but

did not cause boundary layer reattachment. This result contrasts

with earlier work on less aggressive airfoils e.g., Reimann et al.

25on the Pack B, which showed reattachment for at least part

of the wake passing period or a reduction in bubble size in cases

with already reattached boundary layers e.g., Schobeiri et al.

11. The much thicker separation bubble on the L1A prevents

the mixing associated with the wake from penetrating all the way

to the airfoil surface and causing full reattachment.

In the present study, the effect of wakes is considered with

Re= 25,000 and 50,000 under low 0.6% and high 4%

freestream turbulence conditions. Cases with various wake passing frequencies and flow coefficients are documented. Surface

pressure distributions, total pressure loss profiles, and instantaneous boundary layer velocity measurements are used to explain

the mechanism by which the wakes affect separation.

2 Experimental Facility and Measurements

Experiments were conducted in a closed loop wind tunnel with

a seven blade linear cascade, as shown in Fig. 1. A fine screen

located upstream of the cascade is used to break up the boundary

layers that form upstream of the test section and to provide uniform inlet conditions to the cascade. The freestream turbulence

entering the cascade was measured with a cross-wire probe positioned just upstream of the center blade. The turbulence intensities

are 0.8% in the streamwise component and 0.5% in the cross

stream components. The integral length scale of the streamwise

component is 6.3 cm. To produce high freestream turbulence, the

screen is replaced with a coarse grid, consisting of a 1.5 mm thick

sheet metal plate with 19 mm square holes spaced 25.4 mm apart,

center to center, in both directions. In a plane perpendicular to the

inlet flow and 1.7Cx upstream of the center blade, the grid produced uniform flow with TI= 6.0% in the streamwise component

and 4.2% in the cross stream components for an overall intensity

of 4.9%. The streamwise component was also measured at the

inlet plane of the cascade in the four center passages, where it had

decayed to between 4% and 4.2% between blades B2 and B5 and

4.6% between blades B5 and B6. Downstream of the cascade, the

local TI is 1.8% across all passages, as documented in Volino et

al. 22. The local freestream turbulence intensity in the passage at

the beginning of the adverse pressure gradient region is 1.4%. The

change in TI through the passage is due mainly to the change in

the local freestream velocity along with some decay of the turbulence. The upstream integral length scales of the freestream turbulence are 0.12Cx in the streamwise component and 0.04Cx in

the other components. Further details of the facility and inlet flow

are in Volino et al. 22.

A tailboard, shown in Fig. 1, is needed to produce the correct

exit flow angle from the cascade. Its position was set to produce

periodicity at high Reynolds numbers. A tailboard on the opposite

side of the cascade and inlet guide vanes were found to be unnecessary. To produce the correct approach flow to the end blades B1

and B7, the amount of flow escaping around the two ends of the

cascade is controlled with the flaps shown in Fig. 1. The inlet flow

angle was checked with a three-hole pressure probe and found to

agree with the design angle to within 2 deg of uncertainty. Good

periodicity at high Reynolds numbers was obtained in the exit

flow. At low Reynolds numbers, when significant separation

bubbles are present, the periodicity is not as good due to suppression of the separation bubble thickness on the blades closest to the

tailboard. In cases where wakes or other flow control suppress

separation, periodicity is reestablished. The lack of periodicity in

cases with large separation bubbles is considered acceptable since

the focus of the research is separation control, and not documentation of cases with large separation that would be unacceptable in

practice. This compromise facilitates the study of a larger number

of cases by obviating the need to adjust the tailboard by trial and

error for each case. It also provides for better repeatability in the

experiments since the position of the tailboard is fixed for all

cases. Any changes in separation with wakes will be larger in

practice than documented in the experiment due to the effect of

the tailboard in suppressing the bubble size in the no-wake cases.

The wake generator includes a chain near each endwall of the

cascade that passes 0.54Cx upstream of the leading edges of the

cascade blades. The chains then pass downstream around blade

B7 on the inside turn of the cascade and pass well downstream of

the cascade before returning upstream around blade B1 on the

outside turn of the cascade. This completes the chain circuit. A

traverse for probe movement is located within the chain circuit

downstream of the blade row. Each chain is driven by a drive gear

and also passes around six idler sprockets. One of the idler

sprockets is adjustable to maintain tension in the chain. The drive

Table 1 Cascade parameters

Axial

chord, Cx

mm

True

chord

mm

Pitch,

L

mm

Span

mm

Suction

side, Ls

mm

Inlet

flow

angle

deg

Exit

flow

angle

deg

134 146 136 724 203 35 60

tailboard

flap

flap

grid

B1

B2

B3

B4

B5

B6

B7

Fig. 1 Schematic of linear cascade

011011-2 / Vol. 134, JANUARY 2012 Transactions of the ASME

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NASA/CR—2012-217415 190
gears for the upper and lower chains are on a common axle and

driven by a single electric motor so both chains move in unison.

The motor speed is set with a variable frequency inverter. The

chain links have hollow pins, through which the wake generator

rods are attached. Each rod consists of a 4 mm diameter carbon

fiber tube with a steel pin attached at each end. The steel pins are

inserted through the holes in the upper and lower chains, and then

secured with small clips. Tests were run with average distances

between rods of 136 mm, 221 mm, and 272 mm, which correspond to 1L, 1.6L, and 2L, where L is the blade spacing in

the cascade. These ratios of rod to blade spacing are in the range

expected for vane to rotor blade spacing in an engine.

The ratio of the rod diameter to the axial chord is 0.03, which is

consistent with the wake generators of Bons et al. 9and

Funazaki et al. 16. The rods are smaller than those of Kaszeta et

al. 26who had a diameter to chord ratio of 0.06. The present

rods are larger than those of Schobeiri et al. 11and Zhang and

Hodson 15who had rod diameter to chord ratios of about 0.01.

In the present study, as in Bons et al. 9and Kaszeta et al. 26,

the rod wakes are intended to simulate the wakes of very highly

loaded airfoils under low Reynolds number conditions with thick

boundary layers and in some cases large separation bubbles. A

large diameter rod is therefore needed to simulate an airfoil wake

with a large velocity deficit. The studies using smaller diameter

rods were done at higher Reynolds numbers, so separation

bubbles were smaller and rods producing lower velocity deficits

were more appropriate.

For most tests, the rods were driven at a velocity of 1.18 times

the cascade inlet velocity, Ui. This gives a flow coefficient 

=Ui cosi/Urod= 0.7, where i is the inlet flow angle. This is in

the expected range for an engine. The timing of the wake generator is recorded with an infrared photodetector, which senses the

passage of each rod and emits a voltage that is recorded simultaneously as other data are acquired, allowing phase averaging of

the results.

2.1 Measurements. The center blade, designated B4 in Fig.

1, contains pressure taps near the spanwise centerline. Pressure

surveys are made using a pressure transducer 0–870 Pa range

Validyne transducer. Stagnation pressure is measured with a pitot

tube upstream of the cascade and wake generator. The uncertainty

in the suction side pressure coefficients, Cp, is 0.07. Most of this

uncertainty is due to bias error. Stochastic error is minimized by

averaging pressure transducer readings over a 10 s period.

Total pressure losses are documented using a Kiel probe traversed across three blade spacings, 0.63Cx downstream of the

cascade. A traverse is located in the wind tunnel downstream of

the cascade to move the probe. The traverse causes an acceptably

low blockage when it is located at least two Cx downstream of the

cascade.

Pressure and loss surveys were acquired at nominal Re

= 25,000 and 50,000. The Reynolds number, as defined above, is

based on the suction surface length and the nominal cascade exit

velocity. The corresponding Reynolds numbers based on the cascade inlet velocity and the axial chord length are 10,000 and

20,000.

Velocity profiles on the suction surface were measured at the

six streamwise stations listed in Table 2. All stations are downstream of the inviscid pressure minimum at s/Ls= 0.49. Profiles

were acquired near the spanwise centerline of the airfoil with a

hot-wire anemometer AA Lab Systems model AN-1003 and a

single-sensor hot-film probe TSI model 1201-20. The sensor diameter is 51 m, and the active length is 1.02 mm. At each

measurement location, data were acquired for 26 s at a 20 kHz

sampling rate 219 samples. All raw data were saved. The high

sampling rate provides an essentially continuous signal, and the

long sampling time results in low uncertainty in both statistical

and spectral quantities. Data were acquired at 40 wall normal

locations in each profile, extending from the wall to the

freestream, with most points concentrated in the near-wall region.

The probe was positioned as close to tangent to the airfoil surface

as possible at each station, such that the probe body extended

downstream of the sensor and the direction of the traverse was

within 5 deg of normal to the surface. In most cases the closest

point to the wall in each profile was within about 0.2 mm of the

wall, which compares to boundary layer thicknesses ranging from

1.1 mm to over 40 mm.

Flow direction in a separation bubble cannot be determined

with a single-sensor hot-wire, but velocity magnitude can be measured and was found to be near zero within the bubbles of the

present cases when the flow was laminar. In cases where the flow

became turbulent but remained separated, fluctuating velocities

caused false high mean velocity readings in the separation bubble.

With the exception of these turbulent separated cases, the uncertainty in the mean velocity is 3–5% except in the very near-wall

region, where near-wall corrections Wills 27 were applied to

the mean velocity.

Velocity was also measured downstream of the cascade along

the same line used for the total pressure loss measurements.

−0.5 −0.4 −0.3 −0.2 −0.1 0 0.1 0.2 0.3 0.4 0.5 0.5

0.6

0.7

0.8

0.9

1

1.1

1.2

1.3

φ/Lφ

U U/

mean

Rod wake

Cascade wake, Re=50,000, 0.3Cx

Cascade wake, Re=50,000, 0.63Cx

Cascade wake, Re=200,000, 0.3Cx

Cascade wake, Re=200,000, 0.63Cx

−0.5 −0.4 −0.3 −0.2 −0.1 0 0.1 0.2 0.3 0.4 0.5 0

0.05

0.1

0.15

0.2

0.25

φ/Lφ

u′ U/

mean

Rod wake

Cascade wake, Re=50,000, 0.3Cx

Cascade wake, Re=50,000, 0.63Cx

Cascade wake, Re=200,000, 0.3Cx

Cascade wake, Re=200,000, 0.63Cx

Fig. 2 Velocity in wakes of wake generator rods and cascade

wakes with low TI: „a... mean and „b... rms

Table 2 Velocity profile measurement stations

Station 123456

s/Ls 0.53 0.59 0.69 0.78 0.88 0.97

x/Cx 0.65 0.72 0.80 0.86 0.92 0.97

Journal of Turbomachinery JANUARY 2012, Vol. 134 / 011011-3

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NASA/CR—2012-217415 191
Downstream and boundary layer velocity data were both time

averaged and ensemble averaged based on the phase within the

wake passing period. Phase averages of mean and fluctuating velocity are shown below at 24 dimensionless times, t/T, within the

wake passing period, where t is time and T is the period between

wakes.

3 Results

3.1 Wake Characteristics. Several studies have compared

rod wakes to airfoil wakes. Schobeiri et al. 28discussed the

theory for describing wake flows. Pfeil and Herbst 29showed

that the far wake of an airfoil has almost the same wake as a

cylinder having the same drag, although the rods have no means

of matching the correct potential field of adjacent blade rows

30,31. Pluim et al. 10tested rods of various shapes and provided references to earlier studies. They found that the wakes

from wedge shaped rods agreed best with L1A airfoil wakes, but

cylindrical rods also provided a reasonable approximation. Figure

2 shows the mean and rms fluctuating streamwise velocity in the

wakes of the moving rods and the cascade blades of the present

study. The cascade wakes were measured by traversing a hot-wire

probe in lines 0.3Cx and 0.63Cx downstream of the trailing edges

of the blades perpendicular to the axial direction. The variable 

denotes the distance across the cascade. The origin, = 0, is directly downstream of the center blade B4 in the design flow

direction, while /L=−1 and 1 are downstream of blades B3

and B5, respectively. The spacing between airfoil rows in an engine is expected to be about 0.3Cx according to Pluim et al. 10,

so this location would be where the wake from an upstream airfoil

would reach the leading edge of the next row. The moving rod

wakes were measured by placing a stationary hot-wire midway

between blades B4 and B5 in the plane of the leading edges. The

moving rods traversed 0.54Cx upstream of the hot-wire. The data

from the probe were phase averaged on the rod passing cycle to

determine the mean and fluctuating velocity as functions of time.

For comparison in the figures, the rod velocity was then used to

convert time to distance traversed. Curves in the figures were

shifted along the horizontal axis to align the peaks for comparison. All curves were normalized using the average local mean

velocity.

The wakes from the upstream rods, when shown in dimensionless form as in Fig. 2, did not depend on Reynolds number. For all

of the flow and rod velocities considered, the Reynolds number

based on rod diameter is between about 400 and 2000, which

should result in laminar separation for the boundary layer on the

rod and similar wakes for all cases. The wake for low TI, Re

= 50,000, = 0.7, 1.6L rod spacing, and dimensionless wake passing frequency of F= 0.28 is shown. In contrast to the rod wakes,

the wakes of the cascade airfoils depend strongly on Reynolds

number, as shown in Volino 20. Due to a large separation

bubble, the airfoil wake at Re= 50,000 has a large mean velocity

deficit at 0.3Cx, which decreases 80% in peak magnitude by

0.63Cx. The peak turbulence intensity drops by 40% over this

−0.5 −0.4 −0.3 −0.2 −0.1 0 0.1 0.2 0.3 0.4 0.5 0.7

0.75

0.8

0.85

0.9

0.95

1

1.05

1.1

1.15

φ/Lφ

U/U

mean

Rod wake

Cascade wake

−0.5 −0.4 −0.3 −0.2 −0.1 0 0.1 0.2 0.3 0.4 0.5 0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

0.2

φ/Lφ

u′/U

mean

Rod wake

Cascade wake

Fig. 3 Velocity in wakes of wake generator rods and phase

averaged velocity in wakes of cascade blades for low TI, Re

=50,000 case with VGJ flow control †24‡: „a... mean and „b... rms

−1.5

−1

−0.5

0

0.5

1

1.5

10−2

10−1

100

101

102

0

1

2

3

4

5

6

7

x 10−3

φ/Lφ

F

f⋅PSD/U2

e

0

0.2

0.4

0.6

0.8

1

1.2

1.4

10−2

10−1

100

101

102

0

1

2

3

4

5

x 10−3

φ/Lφ

F

f⋅PSD/U2

e

Fig. 4 Wavelet spectra of wake velocity: „a... cascade blades

for low TI, Re=50,000 case with VGJ flow control †24‡ and „b...

wake generator rods

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NASA/CR—2012-217415 192
distance. At Re= 200,000 the boundary layer remains attached, so

the airfoil wake is about half as wide as at the lower Re, and the

peak velocity deficit is about 25% as large. The peak turbulence

intensity is 7%, which is about a third that at the lower Re. The

change in the wake between 0.3Cx and 0.63Cx is much smaller

than at the lower Re. The rod wake lies between the high and low

Re airfoil wakes. Its peak turbulence intensity is 14%. The velocity deficit and turbulence level in the rod wakes in Fig. 2 appear to

be reasonable approximations for an airfoil wake, although they

do not fully match either the high or low Re case. One could

conceivably use a rod of either larger or smaller diameter to match

the low or high Re airfoil cases better. In cases with high

freestream turbulence, the rod wakes are nearly the same as those

in Fig. 2, with 14% peak turbulence intensity. The only difference

is between the wakes, where the background turbulence rises to

4% with high TI.

The cases of the present study will be used for comparison to

future cases with wakes and vortex generator jet flow control.

Figure 3 compares the upstream rod wake to the airfoil wake in a

case with Re= 50,000 and pulsed VGJ blowing with F

= 0.56, 10% duty cycle and blowing ratio jet velocity/freestream

velocity equal to 1. The VGJs provided good flow control in this

case. The boundary layer separates very briefly but reattaches.

Further details are available in Volino et al. 24. For the airfoil

wake in Fig. 3, the phase averaged results at 24 phases through the

VGJ pulsing cycle are shown. Since the boundary layer approaches separation during part of the cycle, the wake velocity

profile magnitude and position change somewhat during the cycle.

The mean velocity deficit and wake width of the upstream rod and

airfoil are about the same. The peak turbulence intensity in the rod

wake is about 40% higher than the highest value in the airfoil

wake.

Turbulence spectra provide additional information about the

wakes. Spectra of the fluctuating streamwise velocity were computed for the cascade airfoil wakes and the rod wakes. Without

flow control, the airfoil wakes at Re= 50,000 exhibit sharp, high

amplitude, low frequency F= 0.7 peaks associated with the shedding frequency of the separation bubble. With Re= 200,000, the

amplitude of the peaks is about two orders of magnitude lower

than at the lower Re, and the peaks are broad banded and centered

at F= 2.5. These peaks are likely associated with the turbulence in

the airfoil boundary layers. The spectra of the upstream rod wakes

have amplitude and frequency between the high and low Re cases,

agreeing most closely with the Re= 50,000 VGJ controlled case

mentioned above. Figure 4 shows wavelet spectra computed as

described in Volino 32using the Mexican Hat wavelet for the

airfoil and rod wakes. Power spectral density premultiplied by

frequency is shown as a function of dimensionless frequency and

distance across the cascade. One representative phase within the

VGJ pulsing cycle is shown for the airfoil wake. As in Figs. 2 and

3, time is converted to position for the rod wake. For the airfoil

wake there is a large peak at F= 1.8 with a magnitude of 0.0052

and a smaller peak at F= 0.6 with a magnitude of 0.0029 that is

hidden behind the larger peak in Fig. 4a. The magnitude of the

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No wakes

2Lφ, ζ=0.7

1.6Lφ, ζ=0.7

1Lφ, ζ=0.7

1.6Lφ, ζ=0.35

Inviscid

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4

B5 B6

No wakes

2Lφ, ζ=0.7

1.6Lφ, ζ=0.7

1Lφ, ζ=0.7

1.6Lφ, ζ=0.35

Fig. 5 Pressure profiles for low TI, Re=25,000 cases: „a... Cp

and „b... total pressure loss

φ/Lφ

t T/ −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

φ/Lφ

t T/ −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Fig. 6 Time-space plot of phase averaged velocity 0.63Cx

downstream of cascade for low TI, Re=25,000, =0.7, and 2L

rod spacing: „a... U/Ue and „b... u/Ue

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NASA/CR—2012-217415 193
two peaks varies during the VGJ cycle. The rod wake has a large

peak at F= 0.8 with a magnitude of 0.0047 and a smaller peak at

F= 0.2 with a magnitude of 0.0016. Figures 2–4 show that the

velocity deficit, turbulence intensity, and spectral content are of

the same order for the airfoil and rod wakes, although the agreement with any one case is not perfect. Since the wake of an

upstream airfoil will not necessarily match exactly the wake of the

airfoil on which it impinges, the present rod wakes are considered

to provide a reasonable approximation to the wakes that might be

shed from an upstream blade row at low Reynolds numbers.

3.2 Low TI

3.2.1 Re= 25,000. The Cp and total pressure loss profiles for

cases with Re= 25,000 and low TI are shown in Fig. 5. The inviscid Cp profile for the L1A airfoil is shown for comparison. The

low peak followed by a plateau in the case without wakes indicates separation without reattachment. For the cases with = 0.7,

decreasing the rod spacing from 2L to 1L which corresponds to

increasing the wake passing frequency F from 0.22 to 0.45

causes Cp to approach the inviscid line and reduces the loss by

over 40% compared with the no-wake case. The wakes also cause

the loss peaks to shift to the right in Fig. 4b, indicating greater

flow turning resulting from a thinner separation bubble. The

greater turning agrees with a 20% increase in lift indicated by the

Cp profiles. Also shown in Fig. 5 is a case with 1.6L rod spacing

and the rod speed and F doubled so = 0.35. The higher wake

passing frequency helps to reduce the separation.

Figures 6–9 show time-space plots of the phase averaged mean

and fluctuating velocity 0.63Cx downstream of the blade row for

the cases of Fig. 5. The contours are normalized by the exit velocity Ue. In Fig. 6, the rod spacing is 2L, so the flow in alternating passages is in phase, with the passages between a half cycle

out of phase. The turbulence Fig. 6b shows vertical strips of

moderately high fluctuations at /L=−1.5, 0.5, 0.5, and 1.5

associated with the airfoil wakes. These positions correspond to

the loss peaks in Fig. 5b. Between these peaks the fluctuation

levels range from very low corresponding to the low background

TI to high when the rod wakes pass. The rod wakes proceed at an

angle in the figure, rising from left to right, as they move forward

in time and transit across the cascade. The highest turbulence

peaks occur where the rod wakes interact with the separation

bubble and airfoil wakes. At the position and time of each rod

wake peak in the turbulence is a corresponding velocity deficit in

the mean velocity Fig. 6a. Following the velocity deficit is a

mean velocity maximum as the flow accelerates after the wake

passes. In Fig. 7, the rod spacing is reduced to 1.6L. All passages

are out of phase with each other, and the peaks are stretched along

the time axis compared with Fig. 6 since the rod speed is the same

but the wake passing period is shorter. In Fig. 8, the rod and blade

spacing are equal, so all passages are in phase with each other.

The wake passing frequency is high enough that there is little time

for the boundary layer to separate between wakes. This reduces

φ/Lφ

t T/ −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

φ/Lφ

t T/ −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Fig. 7 Time-space plot of phase averaged velocity 0.63Cx

downstream of cascade for low TI, Re=25,000, =0.7, and 1.6L

rod spacing: „a... U/Ue and „b... u/Ue

φ/Lφ

t T/ −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

φ/Lφ

t T/ −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Fig. 8 Time-space plot of phase averaged velocity 0.63Cx

downstream of cascade for low TI, Re=25,000, =0.7, and 1L

rod spacing: „a... U/Ue and „b... u/Ue

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NASA/CR—2012-217415 194
the peak turbulence level. There is still a mean velocity deficit

associated with each rod wake followed by a velocity rise, but the

difference between the minima and maxima is reduced since suppression of the separation bubble reduces the flow unsteadiness.

Figure 9 shows the effect of increasing the rod speed. Again, the

increased wake passing frequency helps to suppress the separation

and reduce unsteadiness.

Figure 10 shows time averaged mean and fluctuation velocity

profiles at the six streamwise stations of Table 2. Without wakes

the separation bubble is very thick, and the peak fluctuating velocity is in the shear layer over the bubble. As shown in Volino

20, the shear layer does transition to turbulent between stations 3

and 4, but this does not cause reattachment. Wakes with rod spacing of 2L and 1.6L reduce the separation bubble thickness by

about half, but do not eliminate it. Figure 11 shows phase averaged mean velocity for the 2L spacing case. The six columns

correspond to the six streamwise stations, and the rows correspond to different phases in the wake passing cycle. The boundary

layer separates with a very thin bubble at station 2 that becomes

thicker at the downstream stations at most phases. The bubble

appears as a region of nearly constant, but nonzero velocity near

the wall. The nonzero velocity is an artifact of the hot-wire sensor’s inability to distinguish flow direction and the turbulent fluctuations within the bubble. There is also a period of attached flow

at each station where the velocity far from the wall is somewhat

lower than at other phases and the near-wall velocity more

smoothly approaches zero, as fluid from the outer flow fills in the

region near the wall when the separation bubble collapses. The

reattachment occurs at t/T= 0.125− 0.25, t/T= 0.167− 0.333, t/T

= 0.25− 0.458, t/T= 0.375− 0.583, and t/T= 0.458− 0.708 at stations 2–6, respectively.

More results for the low TI, Re= 25,000 cases are shown in Fig.

12 as contour plots of the phase averaged velocity at different

phases within the wake passing cycle. The vertical axis shows

distance from the wall and the horizontal axis shows distance

along the suction surface. The first and third rows in each case

show mean velocity, and the second and fourth rows show the

corresponding fluctuating velocity. Figure 12a shows the same

case as Fig. 11. The effect of the rod wake appears in the mean

velocity as elevated freestream velocity at t/T= 0.958. The location of high freestream velocity is farther downstream at t/T

= 0.042 and reaches the trailing edge at about t/T= 0.292. During

φ/Lφ

t T/ −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

φ/Lφ

t T/ −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Fig. 9 Time-space plot of phase averaged velocity 0.63Cx

downstream of cascade for low TI, Re=25,000, =0.35, and

1.6Lrod spacing: „a... U/Ue and „b... u/Ue

0 1 0 1 0 1 0 1 0 1 0 1 0

0.1

0.2

U/Ue

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0

0.1

0.2

u′/Ue

y/Ls

No wakes 2Lφ 1.6Lφ

Fig. 10 Time averaged velocity profiles at six streamwise stations for low TI, Re=25,000, and =0.7 cases: top—mean and

bottom—rms

0

0.1

0.083

t/T St. 12 3 4 5 6

0

0.1

0.167

0

0.1

0.25

0

0.1

0.333

0

0.1

0.417

0

0.1

0.5

y/Ls

0

0.1

0.583

0

0.1

0.667

0

0.1

0.75

0

0.1

0.833

0

0.1

0.917

0 1 0 1 0 1 0 1 0 1 0 1 0

0.1

1

U/Ue

Fig. 11 Phase averaged mean velocity profiles for low TI, Re

=25,000, =0.7, and 2Lrod spacing; columns for six streamwise stations; rows for phases in wake cycle

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NASA/CR—2012-217415 195
this time the boundary layer remains thick, as shown by the dark

blue color in the figure, but the near-wall velocity is slightly

elevated lighter shade. The higher near-wall velocity corresponds to the attached profiles at these phases in Fig. 11. As the

wake passes, the fluctuating velocity rises in the freestream and

near the wall. Between t/T= 0.375 and 0.625 the boundary layer

relaxes to an undisturbed, separated state, and then continues in

this state until the next wake arrives. Figure 12b shows results

with the wake spacing reduced to 1.6L. The wake passing is

similar to that in Fig. 12a, but the dimensionless duration of the

wake passing is longer since the wake travels at the same velocity

but the period is shorter. This results in less time for recovery

between wakes.

Figure 13 presents another view of the wake passing effect as

time-space plots of momentum thickness. The horizontal axis

shows distance along the suction surface and the vertical axis

shows dimensionless time. The data are repeated for two cycles to

show the periodicity in time. The white lines in the figure show

the approximate location of the wake and the calmed region following the wake. The leading edge of the wake affected region in

the boundary layer travels at about the freestream velocity, while

the trailing edges of the wake and calmed regions move at about

50% and 30% of the freestream velocity, respectively. These convection velocities agree with values in the literature, such as those

reported by Stieger and Hodson 33and Zhang and Hodson 15.

Momentum thickness is shown since it is proportional to losses in

the boundary layer. Plots of other quantities such as the shape

factor also show the wake and calmed regions, as shown in

Stieger and Hodson 33. In Fig. 13a, the wake causes a strip of

elevated momentum thickness corresponding to the reattachment

and elevated near-wall turbulence in Figs. 11 and 12a. The periods of high momentum thickness and turbulence produce the

high turbulence peaks observed in the airfoil wake in Fig. 6. In

Fig. 13, during the early part of the calmed period the momentum

thickness is slightly suppressed compared with the value between

wake events. During the calmed period the disturbance has already passed, as shown in Fig. 12a, but the boundary layer remains attached, as shown in Fig. 11. Figure 13b shows results

with 1.6L wake spacing. As in Fig. 13a, the momentum thickness increases as the wake passes and is reduced in the calm

region. With less time for recovery between wakes, the variation

in momentum thickness is lower in Fig. 13b.

The present results are similar to those in other studies with

wakes, such as Schobeiri et al. 11and Zhang and Hodson 15.

In those studies the boundary layer reattached even without

wakes, and the passing wake caused a reduction in the separation

bubble size. In the present study, because the Reynolds numbers

are lower and the pressure gradients are stronger, the boundary

layer only intermittently reattaches as the wake passes. There are

y L/ s 0.05

0.1

0.15

0.2

y L/ s

t/T=0.042

0.05

0.1

0.15

0.2

t/T=0.125 t/T=0.208 t/T=0.292

y/Ls 0.05

0.1

0.15

0.2

y L/ s

s/Ls

t/T=0.375

0.6 0.8

0.05

0.1

0.15

0.2

s/Ls

t/T=0.625

0.6 0.8 s/Ls

t/T=0.875

0.6 0.8 s/Ls

t/T=0.958

0.6 0.8

y L/ s 0.05

0.1

0.15

0.2

y L/ s

t/T=0.042

0.05

0.1

0.15

0.2

t/T=0.208 t/T=0.292 t/T=0.375

y L/ s 0.05

0.1

0.15

0.2

y L/ s

s/Ls

t/T=0.458

0.6 0.8

0.05

0.1

0.15

0.2

s/Ls

t/T=0.542

0.6 0.8 s/Ls

t/T=0.625

0.6 0.8 s/Ls

t/T=0.833

0.6 0.8

U/Ue

0 0.5 1 1.5

u′/Ue

0 0.1 0.2 0.3

Fig. 12 Phase averaged mean „rows 1 and 3... and rms streamwise fluctuating „rows 2 and 4... velocity for low TI, Re=25,000,

and =0.7: „a... 2Lrod spacing and „b... 1.6Lrod spacing

s/Ls

t T/

0.55 0.6 0.65 0.7 0.75 0.8 0.85 0.9 0.95

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

1

2

3

4

5

6

7

8

9

s/Ls

t T/

0.6 0.65 0.7 0.75 0.8 0.85 0.9 0.95

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

0

1

2

3

4

5

6

7

Fig. 13 Phase averaged momentum thickness for low TI, Re

=25,000, and =0.7: „a... 2Lrod spacing and „b... 1.6Lrod spacing; lines show leading edge „solid... and trailing edge „dasheddotted... of wake, and trailing edge of the calm region „dashed...

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NASA/CR—2012-217415 196
several mechanisms by which wakes could induce reattachment.

The elevated freestream turbulence in the wake will cause transition to move upstream in the shear layer, and if transition occurs

before the separation bubble becomes too thick, the turbulent mixing could extend close enough to the wall to cause reattachment.

Volino et al. 22showed, however, that elevated background turbulence alone did not cause reattachment on the L1A at Re

= 25,000. The upstream turbulence intensity in the wake is 14%,

as documented above, but as the flow is accelerated through the

passage, this local intensity drops to about 3% at the beginning of

the adverse pressure gradient region. Separated flow transition

correlations, such as those of Roberts and Yaras 34and Volino

and Bohl 35, indicate that a rise to 3% TI would cause transition

to shift upstream by about 0.06Ls, which may not be enough to

explain the full wake effect on reattachment. Other effects of

wakes are explained in detail by Stieger and Hodson 33and

Zhang and Hodson 15, who used instantaneous velocity field

measurements to document negative jets resulting from temporal

freestream acceleration and rollup vortices in the shear layer that

caused reattachment. These are likely the same mechanisms responsible for reattachment in the present study.

3.2.2 Re= 50,000. Pressure results for low TI cases with Re

= 50,000 are shown in Fig. 14. Without wakes the shear layer does

not reattach, as in the Re= 25,000 case. With wakes and = 0.7,

reattachment is clear for all rod spacings and losses are reduced

by over 40%. With the rod speed reduced to = 1.4 F reduced to

0.14, only partial reattachment is observed. Wake velocity is

shown in Figs. 15–17 for the = 0.7 cases with rod spacings of

2L, 1.6L, and 1L, respectively. The airfoil wakes, appearing as

vertical strips of low mean velocity and about 7% turbulence, are

in the expected positions based on the loss peaks of Fig. 14b.

There is some rise in the magnitude of the turbulence to about

10% when the airfoil and rod wakes combine in each cycle, particularly in the cases with farther rod spacing. Between the airfoil

wakes, the rod wakes appear in the freestream once per cycle with

about 5% turbulence. The mean velocity drops during the rod

wakes and then rises after the wake passes. The variability in the

velocity and turbulence is much lower than in the Re= 25,000

cases because the wakes suppress the separation bubble more at

the higher Re. This is shown in the time averaged velocity profiles

in Fig. 18. Without wakes there is a large separation bubble. With

wakes the boundary layer is attached. The phase averaged mean

velocity profiles are shown in Fig. 19 for the = 0.7, 2L F

= 0.22 case. The wakes affect the freestream velocity, particularly

at the upstream stations, and a thin separation bubble is visible at

stations 3 and 4 between wakes. At the downstream stations the

boundary layer appears attached at all times. Figure 20a shows

contours of the phase averaged mean and fluctuating velocity for

the = 0.7, 2L case. The velocity variation in the freestream is

visible between t/T= 0.042 and 0.375 as the wake passes, and the

freestream turbulence is slightly elevated during this period as

well. The boundary layer turbulence also rises as the wake passes.

The variation during the cycle is not as great as in the Re

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No wakes

2Lφ, ζ=0.7

1.6Lφ, ζ=0.7

1Lφ, ζ=0.7

1.6Lφ, ζ=1.4

Inviscid

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4

B5 B6

No wakes

2Lφ, ζ=0.7

1.6Lφ, ζ=0.7

1Lφ, ζ=0.7

1.6Lφ, ζ=1.4

Fig. 14 Pressure profiles for low TI and Re=50,000 cases: „a...

Cp and „b... total pressure loss

φ/Lφ

t T/ −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

φ/Lφ

t T/ −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Fig. 15 Time-space plot of phase averaged velocity 0.63Cx

downstream of cascade for low TI, Re=50,000, =0.7, and 2L

rod spacing: „a... U/Ue and „b... u/Ue

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NASA/CR—2012-217415 197
= 25,000 cases of Fig. 12. The boundary layer remains relatively

thin and attached at all phases. The time-space plot of the momentum thickness is shown for this case in Fig. 21a. As in the lower

Re cases, the wake causes an increase in momentum thickness as

it passes. Any delay in momentum thickness growth in the calmed

region is slight. The overall value of the momentum thickness is

less than half the value in the Re= 25,000 cases, which corresponds to the lower losses in Fig. 14 compared with Fig. 5 for the

= 0.7, 2L cases and the lower wake turbulence in Fig. 15 compared with Fig. 6. A case was also documented with = 0.7 and

1.6L. It showed little difference from the = 0.7, 2L case. A

reduction of the wake rod velocity to = 1.4, 1.6L F= 0.14 produces the results in Figs. 20b and 21b. With the reduced wake

frequency the boundary layer is separated for much of the cycle

and the momentum thickness is over double its value in the 

= 0.7 case. These results agree with the high loss and separated

flow Cp profile with = 1.4 in Fig. 14.

3.3 High TI. The pressure results for the high freestream turbulence cases with Re= 25,000 are shown in Fig. 22. Wake velocity contours are shown in Figs. 23 and 24. The results are similar

to the corresponding low TI results of Figs. 5, 7, and 8. The Cp

distributions indicate slightly less separation with high TI, and in

agreement the mean velocities show less variation and the turbulence values are slightly lower with high TI.

Results for the Re= 50,000 cases are shown in Figs. 25–27. For

the = 0.7 cases, the results are nearly the same as in the corresponding low TI cases of Figs. 14, 16, and 17. In the 1.6L, 

= 1.4 cases separation and losses are higher than with lower , but

are clearly reduced by the high TI. The effect of the rod wakes

φ/Lφ

t T/ −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

φ/Lφ

t T/ −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Fig. 16 Time-space plot of phase averaged velocity 0.63Cx

downstream of cascade for low TI, Re=50,000, =0.7, and 1.6L

rod spacing: „a... U/Ue and „b... u/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

φ/Lφ

t T/ −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Fig. 17 Time-space plot of phase averaged velocity 0.63Cx

downstream of cascade for low TI, Re=50,000, =0.7, and 1L

rod spacing: „a... U/Ue and „b... u/Ue

0 1 0 1 0 1 0 1 0 1 0 1 0

0.1

0.2

U/Ue

y/Ls

St 12 3 4 5 6

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0

0.1

0.2

u′/Ue

y/Ls

No wakes 2Lφ 1.6Lφ

Fig. 18 Time averaged velocity profiles at six streamwise stations for low TI, Re=50,000, and =0.7 cases: top—mean and

bottom—rms

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NASA/CR—2012-217415 198
appears to dominate over the background freestream turbulence

effect, and if the wake passing frequency is high enough to suppress separation, the TI level between wakes is largely irrelevant.

If there is sufficient time between wakes for the boundary layer to

separate, high background TI helps to limit the separation.

In agreement with results from the literature for other airfoils,

wakes help to suppress boundary layer separation. For the L1A

airfoil with Re= 25,000, a dimensionless wake passing frequency

of about F= 0.5 appears sufficient to effectively eliminate separation. At lower frequencies the boundary layer separates between

wakes. For Re= 50,000, F= 0.3 is sufficient to eliminate separation. It does not appear to matter if a particular frequency is

achieved by reducing the spacing between wakes or by increasing

the wake generator velocity, so long as F is sufficiently high. The

required wake frequencies are about equal to the vortex generator

jet pulsing frequencies required for separation control with steady

inflow observed by Volino et al. 24,36.

4 Conclusions

The effect of unsteady wakes on the flow over the very high lift

L1A airfoil was studied experimentally under low and high

freestream turbulence conditions. Reynolds numbers based on

suction surface length and nominal exit velocity of 25,000 and

50,000 were considered. Without wakes, the boundary layer separated and did not reattach. Wakes shed by upstream rods caused

the boundary layer to reattach. When the wake passing frequency

was low, either due to large spacing between wake generator rods

or low rod velocity, the boundary layer separated between wakes.

With a dimensionless wake passing frequency of F= 0.5 or higher

at Re= 25,000 or with F= 0.3 or higher at Re= 50,000, separation

was largely suppressed. Background freestream turbulence also

helps to reduce separation, but the effect is secondary to the wake

effect.

The expected wake passing frequency in an engine is near the

frequency at which separation begins to be a problem on the L1A.

The present cases will serve as baseline results for future cases in

which wakes and flow control with vortex generator jets are combined.

Acknowledgment

This work was sponsored by the National Aeronautics and

Space Administration under Grant No. NNC07IA10I. The grant

monitors were Dr. Anthony Strazisar and Dr. James Heidmann of

the NASA Glenn Research Center. The support of the United

States Naval Academy Technical Support Department Shop and

Fluids Laboratory is greatly appreciated.

Nomenclature

Cp  2PT−P/Ue

2

, pressure coefficient

Cx  axial chord length

F  fLj-te /Uave, dimensionless frequency

f  frequency

Lj-te  length of adverse pressure gradient region on

the suction surface

Ls  suction surface length

L  blade spacing pitch

P  pressure

PS  upstream static pressure

PT  upstream stagnation pressure

0

0.1

0.083

t/T St. 12 3 4 5 6

0

0.1

0.167

0

0.1

0.25

0

0.1

0.333

0

0.1

0.417

0

0.1

0.5

y/Ls

0

0.1

0.583

0

0.1

0.667

0

0.1

0.75

0

0.1

0.833

0

0.1

0.917

0 1 0 1 0 1 0 1 0 1 0 1 0

0.1

1

U/Ue

Fig. 19 Phase averaged mean velocity profiles for low TI, Re

=50,000, =0.7 cases, and 2Lrod spacing; columns for six

streamwise stations; rows for phases in wake cycle

y/Ls 0.05

0.1

0.15

0.2

y/Ls

t/T=0.042

0.05

0.1

0.15

0.2

t/T=0.125 t/T=0.208 t/T=0.292

y L/ s 0.05

0.1

0.15

0.2

y L/ s

s/Ls

t/T=0.375

0.6 0.8

0.05

0.1

0.15

0.2

s/Ls

t/T=0.625

0.6 0.8 s/Ls

t/T=0.875

0.6 0.8 s/Ls

t/T=0.958

0.6 0.8

y L/ s 0.05

0.1

0.15

0.2

y/Ls

t/T=0.042

0.05

0.1

0.15

0.2

t/T=0.208 t/T=0.375 t/T=0.542

y L/ s 0.05

0.1

0.15

0.2

y L/ s

s/Ls

t/T=0.667

0.6 0.8

0.05

0.1

0.15

0.2

s/Ls

t/T=0.833

0.6 0.8 s/Ls

t/T=0.917

0.6 0.8 s/Ls

t/T=1

0.6 0.8

U/Ue

0 0.5 1 1.5

u′/Ue

0 0.1 0.2 0.3

Fig. 20 Phase averaged mean „rows 1 and 3... and rms streamwise fluctuating „rows 2 and 4... velocity for low TI and Re

=50,000: „a... =0.7,2Land „b... =1.4,1.6L

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NASA/CR—2012-217415 199
s/Ls

t/T

0.55 0.6 0.65 0.7 0.75 0.8 0.85 0.9 0.95

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

0.5

1

1.5

2

2.5

3

s/Ls

t T/

0.6 0.65 0.7 0.75 0.8 0.85 0.9 0.95

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

1

2

3

4

5

6

Fig. 21 Phase averaged momentum thickness for low TI and

Re=50,000: „a... =0.7,2Land „b... =1.4,1.6L

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No wakes

1.6Lφ, ζ=0.7

1Lφ, ζ=0.7

Inviscid

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4

B5 B6

No wakes

1.6Lφ, ζ=0.7

1Lφ, ζ=0.7

Fig. 22 Pressure profiles for high TI and Re=25,000 cases: „a...

Cp and „b... total pressure loss

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NASA/CR—2012-217415 200
φ/Lφ

t T/ −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

φ/Lφ

t T/ −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Fig. 23 Time-space plot of phase averaged velocity 0.63Cx

downstream of cascade for high TI, Re=25,000, =0.7, and

1.6Lrod spacing: „a... U/Ue and „b... u/Ue

φ/Lφ

t T/ −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

φ/Lφ

t T/ −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Fig. 24 Time-space plot of phase averaged velocity 0.63Cx

downstream of cascade for high TI, Re=25,000, =0.7, and 1L

rod spacing: „a... U/Ue and „b... u/Ue

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NASA/CR—2012-217415 201
0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

s/Ls

Cp

No wakes

1.6Lφ, ζ=0.7

1Lφ, ζ=0.7

1.6Lφ, ζ=1.4

Inviscid

−1.5 −1 −0.5 0 0.5 1 1.5

0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4

B5 B6

No wakes

1.6Lφ, ζ=0.7

1Lφ, ζ=0.7

1.6Lφ, ζ=1.4

Fig. 25 Pressure profiles for high TI and Re=50,000 cases: „a...

Cp and „b... total pressure loss spacing: „a... U/Ue and „b... u/Ue

φ/Lφ

t T/ −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

φ/Lφ

t T/ −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Fig. 26 Time-space plot of phase averaged velocity 0.63Cx

downstream of cascade for high TI, Re=50,000, =0.7, and

1.6Lrod spacing: „a... U/Ue and „b... u/Ue

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NASA/CR—2012-217415 202
PTe  downstream stagnation pressure

Re  UeLs/	, exit Reynolds number

s  streamwise coordinate, distance from the leading edge

T  period of jet pulsing cycle

t  time

TI  background freestream turbulence intensity

U  local mean velocity

Uave  average freestream velocity in the adverse

pressure gradient region

Ui  inlet freestream velocity

Ue  nominal exit freestream velocity based on inviscid solution

u  rms fluctuating streamwise velocity

x  axial distance from the leading edge

  coordinate along the blade spacing, normal to

the axial chord

	  kinematic viscosity

  density


  PT−PTe/PT−PS, total pressure loss

coefficient

  Ui cosi/Urod=Uaxial/Urod, flow coefficient

References

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φ/Lφ

t T/ −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

φ/Lφ

t T/ −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Fig. 27 Time-space plot of phase averaged velocity 0.63Cx

downstream of cascade for high TI, Re=50,000, =0.7, and 1L

rod spacing: „a... U/Ue and „b... u/Ue

Journal of Turbomachinery JANUARY 2012, Vol. 134 / 011011-15

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NASA/CR—2012-217415 203
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011011-16 / Vol. 134, JANUARY 2012 Transactions of the ASME

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NASA/CR—2012-217415 204
Appendix B 

Additional results 

 

NASA/CR—2012-217415 205
6. Additional Results 

The following figures show the ensemble averaged jet velocity measured at the exit of a VGJ 

hole with the main flow in the wind tunnel off. 

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

0.45

0.5

t/T

B

3 Hz, B=0.5

DC=50%

10%

2%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

t/T

B

3 Hz, B=0.75

DC=50%

10%

2%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

t/T

B

3 Hz, B=1.0

DC=50%

10%

2%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

t/T

B

3 Hz, B=1.5

DC=50%

10%

2%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

t/T

B

3 Hz, B=2.0

DC=50%

10%

2%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

t/T

B

3 Hz, B=2.5

DC=50%

10%

2%

NASA/CR—2012-217415 206
0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

3

t/T

B

3 Hz, B=3.0

DC=50%

10%

2%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.05

0.1

0.15

0.2

0.25

t/T

B

6 Hz, B=0.25

DC=50%

10%

2%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

0.45

0.5

t/T

B

6 Hz, B=0.50

DC=50%

10%

2%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

t/T

B

6 Hz, B=0.75

DC=50%

10%

2%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

t/T

B

6 Hz, B=1.0

DC=50%

10%

2%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

t/T

B

6 Hz, B=1.5

DC=50%

10%

2%

NASA/CR—2012-217415 207
0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

t/T

B

6 Hz, B=2.0

DC=50%

10%

2%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

t/T

B

6 Hz, B=2.5

DC=50%

10%

2%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

3

t/T

B

6 Hz, B=3.0

DC=50%

10%

2%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.05

0.1

0.15

0.2

0.25

t/T

B

12 Hz, B=0.25

DC=50%

10%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

0.45

0.5

t/T

B

12 Hz, B=0.50

DC=50%

10%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

t/T

B

12 Hz, B=0.75

DC=50%

10%

NASA/CR—2012-217415 208
0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

t/T

B

12 Hz, B=1.0

DC=50%

10%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

t/T

B

12 Hz, B=1.5

DC=50%

10%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

t/T

B

12 Hz, B=2.0

DC=50%

10%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

t/T

B

12 Hz, B=2.5

DC=50%

10%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

3

t/T

B

12 Hz, B=3.0

DC=50%

10%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.05

0.1

0.15

0.2

0.25

t/T

B

24 Hz, B=0.25

DC=50%

10%

NASA/CR—2012-217415 209
0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

0.45

0.5

t/T

B

24 Hz, B=0.50

DC=50%

10%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

t/T

B

24 Hz, B=0.75

DC=50%

10%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

t/T

B

24 Hz, B=1.0

DC=50%

10%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

t/T

B

24 Hz, B=1.5

DC=50%

10%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

t/T

B

24 Hz, B=2.0

DC=50%

10%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

t/T

B

24 Hz, B=2.5

DC=50%

10%

NASA/CR—2012-217415 210
0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

3

t/T

B

24 Hz, B=3.0

DC=50%

10%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.05

0.1

0.15

0.2

0.25

t/T

B

48 Hz, B=0.25

DC=50%

10%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

0.45

0.5

t/T

B

48 Hz, B=0.50

DC=50%

10%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

t/T

B

48 Hz, B=0.75

DC=50%

10%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

t/T

B

48 Hz, B=1.0

DC=50%

10%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

t/T

B

48 Hz, B=1.5

DC=50%

10%

NASA/CR—2012-217415 211
0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

t/T

B

48 Hz, B=2.0

DC=50%

10%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

t/T

B

48 Hz, B=2.5

DC=50%

10%

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 0

0.5

1

1.5

2

2.5

3

t/T

B

48 Hz, B=3.0

DC=50%

10%

NASA/CR—2012-217415 212
The following section is a taken from the progress report for January – June 2009. It 

provides additional results and figures in addition to those published in the papers included 

above. 

Summary 

During this period, data were acquired for experimental cases with flow control using Vortex 

Generator Jets (VGJs) under low and high freestream turbulence conditions. Various jet 

amplitude, pulsing frequency and duty cycle combinations were considered. A paper was 

presented at Turbo Expo in Orlando (Volino et al., 2009) showing the pressure profile and total 

pressure loss results for the low freestream turbulence cases. The effect of jet spacing was 

considered by blocking some of the jets. Additionally, the wake generator was completed, and 

experimental cases were documented with unsteady wakes under low freestream turbulence 

conditions. Phase averaging of the velocity data has been done to show the behavior of the 

boundary layer and wake velocity as a function of time during the jet pulsing cycle or the wake 

passing period. Spectral processing using Fourier and wavelet analysis has been done on all the 

velocity data to determine the important frequencies in the fluctuating velocity, again as a 

function of time during the pulsing or wake cycle. 

Computationally, unsteady RANS simulations were completed using the Transition-sst 

model for cases with steady and pulsed VGJs. They did not predict the flow control with VGJs 

observed in the experiments. Large eddy simulations (LES) were more successful, correctly 

predicting which cases had flow reattachment. A conference paper (Ibrahim et al., 2009) has 

been accepted for presentation showing results for cases with steady VGJs. 

The next experiments will be with combined wakes and VGJs. The valves for the pulsing 

will be driven using a signal from the wake generator. A delay can be programmed into the 

circuit to set any delay desired to time the VGJs to the wakes. 

The CFD work continues, looking at unsteady wakes next. In addition, more work will be 

done with unsteady RANS calculations to try and better predict the flow with a finer grid than 

used previously, but still courser than required for LES. 

Facility 

The wind tunnel, cascade and pulsing hardware were described in previous progress reports 

and in Volino (2008), Volino et al. (2008), Ibrahim et al. (2008), and Volino et al. (2009). Since 

then, the wake generator has been completed and used. It includes a chain near each endwall of 

the cascade which passes 0.54 Cx (Cx=axial chord) upstream of the leading edges of the cascade 

blades. The chain then passes downstream around the end blade on the inside turn of the cascade 

and passes well downstream of the cascade before returning upstream around the opposite end 

blade on the outside turn of the cascade. This completes the chain circuit. Each chain is driven 

by a drive gear and also passes around six idler sprockets. One of the idler sprockets is 

adjustable to maintain tension in the chain. The drive gears for the upper and lower chains are 

on a common axle and driven by a single electric motor so both chains move in unison. The 

motor speed is set with a variable frequency inverter. The chain links have hollow pins, through 

which the wake generator rods are attached. Each rod consists of a 4 mm diameter carbon fiber 

tube with a steel pin attached at each end. The steel pins are inserted through the holes in the 

upper and lower chain, and then secured with small clips. The rods were spaced 17 or 18 chain 

links apart, for an average distance between rods of 221 mm. This compares to the blade spacing 

in the cascade of 136 mm. The ratio of rod to blade spacing (1.6) is at the upper end of what is 

NASA/CR—2012-217415 213
expected for vane to rotor spacing in an engine. This spacing was chosen instead of a closer 

spacing because it is more challenging for flow control. 

The rods were driven at a velocity of 1.18 times the cascade inlet velocity, Ui. This gives a 

flow coefficient, Uicos(αi)/Urod = 0.7, where αi is the inlet flow angle. This is in the expected 

range for an engine. For the pulsed VGJ cases, a dimensionless frequency was defined as 

F=fLj-te/Uave, where f is the pulsing frequency in Hz, Lj-te is the distance from the VGJ holes to 

the trailing edge, and Uave is the average freestream velocity between the VGJ holes and the 

trailing edge. Using this same definition for the wake passing, the wake passing frequency for 

the present experiments is F=0.28. This corresponds to one of the VGJ frequencies studied 

previously. For a blade Reynolds number of 50,000 (based on exit velocity and suction surface 

length), the rod velocity is 2.2 m/s. The wake generator was run for an extended period at this 

speed. It was also operated at double this speed, but in these cases a rod broke loose after a short 

time. It is possible that adjustments could be made to allow sustained higher speed operation, 

but this is likely not necessary since the cases of most interest for flow control are at Re=50,000 

and below. In addition to the cases with F=0.28, the wake generator speed was adjusted for a 

few cases with F above or below 0.28 to investigate the effect of wake frequency. 

The timing of the wake generator is recorded with an infrared photo detector located near the 

lower endwall between the cascade leading edge and the wake generator. The detector emits 

light, and when a wake rod passes, it reflects the light back into the detector. The voltage signal 

from the detector is recorded simultaneously as other data are acquired, allowing phase 

averaging locked to the wake passing. The detector signal is also sent to a function generator, 

which produces a square wave for driving the solenoid valves for the VGJs. The duration of the 

high voltage in the square wave and the delay between the detector signal and the function 

generator output can be set to produce any duty cycle desired for the VGJs and any delay desired 

between the wake passing and the VGJ pulse. 

Results 

VGJ Cases, Low Freestream Turbulence 

Pressure (Cp) profiles from the test airfoil and total pressure loss profiles measured 

downstream of the cascade were presented for multiple cases in the previous progress report and 

in Volino et al. (2009). This paper is attached at the end of this report. The VGJs are located at 

the pressure minimum on the suction surface, s/Ls=0.5, where s is the distance from the leading 

edge along the surface and Ls is the suction surface length. The holes are inclined at 30 degrees 

to the surface and 90 degrees to the main flow direction. The hole spacing is 10.7 diameters or 

0.041Cx. As explained in Volino et al. (2009), the blowing ratio, B, of the VGJs is defined as the 

ratio of the maximum velocity in the jet pulse, measured at the jet centerline, divided by the local 

freestream velocity. The duty cycle, D, is the fraction of time the jets are turned on during each 

pulsing cycle. 

Velocity profiles were acquired for several cases at six streamwise positions along the 

suction side of the test blade. Velocity was also measured 0.63Cx downstream of the cascade in 

the same location used for the total pressure loss measurements. Velocity data were acquired 

using a traversing hot-film probe and digitized at a 20 kHz sampling rate. All raw data were 

saved for subsequent phase averaging and spectral analysis. The velocity data help to explain 

how the VGJs control separation and why certain pulsing frequencies and duty cycles work 

NASA/CR—2012-217415 214
better than others. The data are also used for CFD validation. Velocity results are presented 

next. 

Re=25,000. Figure 1 shows the pressure coefficients for cases with Re=25,000 and steady 

VGJs. The pressure coefficient Cp=2(PT-P)/ρUe

2

, where PT is the upstream total pressure, P is 

the static pressure on the airfoil surface, and Ue is the nominal exit velocity from the cascade 

based on the inlet velocity and the design flow angles. The Cp values are shown verses s/Ls, 

where is the distance along the surface from the leading edge, and Ls is the suction surface 

length. In Volino et al. (2009) it was shown that with steady blowing, a jet to freestream velocity 

ratio of B=2.0 is needed for any effect of the VGJs on Cp, and the effect increases at larger B. 

With pulsed jets, some effect on Cp is observed with B=0.5 if F=0.56 or greater. At higher B, 

pulsing with F=0.28 has some effect. Figure 1 shows cases for which velocity profile data are 

also available. If the difference between the suction and pressure side Cp values is integrated in 

the axial direction, the result is proportional to the lift on the airfoil. Figure 1b shows integrated 

results for the cases in Fig. 1a normalized on the integrated Cp for the inviscid solution for the 

passage. Steady blowing with B=3.0 lowers Cp near the trailing edge on the suction surface and 

increases the overall lift by about 10% over the case with no blowing. Pulsed jets with B=1.0 

and F=0.28 are more effective, particularly when the duty cycle is increased from D=10% to 

50%. At higher frequencies the jets are effective even with D=10%. The Cp profiles are 

significantly different than the inviscid solution in all cases, suggesting that separation is not 

completely controlled, even in the best case. 

Figure 2 shows total pressure loss profiles for the cases in Fig. 1. The total pressure loss is 

measured with a Kiel probe traversed along a line 0.63 Cx downstream of the cascade. The loss 

coefficient is defined as Ψ=(PT-PTe)/(PT-Ps), where PTe is the total pressure downstream of the 

cascade, PT is the upstream total pressure, and Ps is the upstream static pressure. The coordinate 

φ is the spanwise position measured perpendicular to the axial direction with φ=0 corresponding 

to the location directly downstream of the center blade in the direction of the design exit flow 

angle. The coordinate φ is normalized by the blade spacing, Lφ. The labels B3, B4, etc., refer to 

the blades in the cascade. The loss peak labeled B4 corresponds to the wake of the center blade. 

Integrated loss results across one Lφ centered on the B4 peak are shown in Fig. 2b. The 

integrated results also include the loss associated with the jets themselves, as explained in Volino 

et al. (2009). Without flow control there is a large separation bubble which does not reattach. 

As explained in Volino et al. (2008), the tailboard on one side of the cascade reduces the bubble 

size on the nearer blade, so the flow is not periodic across the cascade. Steady blowing with 

B=3.0 lowers the loss peaks and reestablishes periodicity, but the integrated loss still rises 

compared to the uncontrolled case due to the loss associated with the jets themselves at this high 

blowing ratio. In agreement with the Cp profiles, pulsing with B=1.0, F=0.28 and D=10% is not 

very effective, but with increased duty cycle or higher frequency the losses are reduced to about 

70% of the baseline value. 

The velocity profile measurements help explain the Cp and pressure loss results. Figure 3 

shows velocity profiles from the boundary layer at six streamwise locations along the suction 

surface for the baseline case without flow control. The stations are spaced along the surface at 

s/Ls= 0.53, 0.59, 0.69, 0.78, 0.88, and 0.97. The VGJ holes are located at the suction peak at 

s/Ls=0.5. The top row in Fig. 3 shows the local mean velocity and the second row show the rms 

fluctuating streamwise velocity u′. The boundary layer has already separated at the first station, 

NASA/CR—2012-217415 215
and the separation bubble grows larger at the downstream stations. The fluctuating velocity has 

its maximum in the shear layer over the separation bubble. 

Figure 4 shows the velocity results with steady blowing and B=3.0. Separation is delayed 

until Station 2, and the separation bubble is clearly thinner than in the baseline case, but the 

separation bubble is still clearly large and reattachment does not occur. The velocity results 

agree with the Cp and pressure loss results of Figs. 1 and 2, which showed that steady VGJs at 

B=3.0 have some effect, but do not appear to fully control the separation. 

Figure 5 shows time averaged results for pulsed VGJs with F=0.28, D=10% and B=1.0. 

There is clearly a large separation, although the thickness of the separation bubble is smaller than 

in the baseline case. Phase averaged results are shown in Fig. 6 for the same case. Figure 6a 

shows the mean velocity. The format is the same as in Fig. 5, but each row shows a different 

phase in the pulsing cycle. The phase t/T=0, where T is the pulsing period, corresponds to the 

opening of the solenoid valves, which provide air to the cavity in the airfoil that supplies the 

VGJs. At Station 3, the profile appears nearly attached at t/T=0.375. As the flow convects 

downstream, the phase of near attachment at Station 4 is centered at t/T=0.5, continuing to 

t/T=0.625 at Station 5 and t/T=0.708 at Station 6. Figure 6b shows u′ results for this case. At 

the times and locations corresponding to near reattachment in Fig. 6a, the fluctuating velocity is 

somewhat lower than at other phases. 

Pulsing with F=0.28 and D=10% appears to induce boundary layer reattachment, but only for 

a short period of time. The duration of the pulse effect is too short and the pulses occur too 

infrequently to cause reattachment for most of the cycle. When the duty cycle is increased to 

D=50%, as shown in Fig. 7, the time averaged flow appears more attached, although the 

boundary layer is still very thick, the mean velocity profile has an inflection point, and the peaks 

in u′ are far from the wall, as in a separated flow case. The phase averaged results in Fig. 8 show 

that the boundary layer alternates between separated and attached states during the pulsing cycle. 

Separation is delayed until Station 2, and at Station 3 there is a thin but clear separation bubble at 

most phases. At t/T=0.333 and at t/T=0.833, however, the boundary layer appears close to 

reattachment. The t/T=0.333 reattachment occurs at about the same time as observed in the 

D=10% case in Fig. 6. The t/T=0.833 reattachment occurs half a cycle later. It appears that 

turning on the VGJ induces reattachment, and turning off the VGJ at the end of the 50% duty 

cycle also induces reattachment. During the blowing between t/T=0.333 and 0.833, the 

boundary layer separates. As the flow convects downstream, the reattachment phases are 

centered at t/T=0.417 and 0.875 at Station 4, 0.583 and 0.042 at Station 5, and 0.708 and 0.208 at 

Station 6. The duration of the reattached flow increases at the downstream stations. By Station 

6, the boundary layer appears to be attached about 70% of the time, and the separation bubble is 

much thinner during the remaining phases than in the cases considered above. It appears that 

turning the VGJ off is also effective for flow control. Increasing the duty cycle to 50% separates 

the turning on and off of the VGJs compared to the D=10% case, resulting in two distinct flow 

control events for each pulsing cycle. Because the boundary layer has less time to separate 

between events, the boundary layer remains attached more of the time and the separation bubble 

thickness is reduced. Increasing the duty cycle from 10 to 50% effectively doubles the flow 

control frequency and results in significantly better separation control. The fluctuating velocities 

shown in Figs. 8b are lower when the flow is attached than when separated. The same 

observation was made in Fig. 6 for the D=10% case. 

Time averaged results for the F=0.56, D=10%, B=1.0 case are shown in Fig. 9. The 

boundary layer is slightly thinner than in the F=0.28, D=50%, B=1.0 case of Fig. 7, but 

NASA/CR—2012-217415 216
otherwise the two cases appear very similar. Phase averaged results are shown in Fig. 10. 

Separation occurs between Stations 2 and 3 at all phases. By Station 4, reattachment has 

occurred between about t/T=0.792 and 0.917. These dimensionless times are roughly double 

those observed at Station 4 in Figs. 6 and 8, where F=0.28. Hence, the dimensional time 

between the opening of the solenoid valve and reattachment at Station 4 is about the same in all 

three cases. This is consistent with the convection time between the VGJ hole and Station 4 

being the same in all three cases, which are at the same Re=25,000. As in the previous cases, u′

is lower when the flow is attached. The duration of the attached flow increases between Stations 

4 and 5, and by Station 6 the flow is attached for about 90% of the cycle. 

The pulsing frequency is increased to F=1.12, D=10%, B=1.0 in Fig. 11, which shows time 

averaged results. The boundary layer is slightly thinner, but otherwise similar to the previous 

case with F=0.56. The phase averaged results in Fig. 12 show that separation occurs at most 

phases at Station 3, but that at about 25% of the phases the boundary layer remains attached. By 

Station 4 the boundary layer has reattached by t/T=0.45, and remains attached until about 

t/T=0.75. The dimensional times corresponding to this period match those observed when 

F=0.28 and F=0.56 in Figs. 6, 8 and 10. As in the previous cases, the fluctuating velocity is 

lower during the attached flow period. The separation bubble remains thin at the other phases. 

By Station 5 the boundary layer is attached 70% of the time. By Station 6 it appears attached for 

the entire cycle, although some profiles have inflection points and the boundary layer is still 

relatively thick compared to cases at higher Reynolds numbers. 

In summary, with low freestream turbulence and Re=25,000 on the present L1A airfoil, a 

disturbance frequency of about 0.56 is needed to provide effective separation control. This can 

be achieved by pulsing with F=0.56 or by pulsing with F=0.28 and a 50% duty cycle so that the 

switching on and off of the jets provide two effective disturbances spaced half a cycle apart for 

every pulse. A moderate blowing ratio of 1.0 is sufficient. When the separation is controlled, 

the lift as indicated by the Cp profiles rises by about 28% compared to the baseline case, and the 

total pressure loss pressure loss drops by about 28%. The switching on and off of the VGJs 

provides a more effective disturbance for flow control than steady blowing. Although blowing 

with steady jets and B=3.0 reduced the thickness of the separation bubble slightly compared to 

the baseline case, the boundary layer did not reattach and losses remained high in this case. The 

adverse pressure gradient is sufficiently strong at this Reynolds number that the boundary layer 

remains close to separation even when the flow control is effective. Hence the boundary layer is 

significantly thicker than observed at higher Reynolds numbers, the Cp profile does not match 

the inviscid solution, and the total pressure losses remain over double those observed at high Re. 

Re=50,000. The Cp profiles for several cases with Re=50,000 and the integrated Cp results 

are shown in Figs. 13. Without flow control, the boundary layer separates and does not reattach, 

as in the Re=25,000 case. Steady jets at B=2.0 control separation and increase the lift by 20%. 

Pulsed jets at F=0.14 with D=10% and B=1.0 have a small effect, but the lift is only increased 

about 8% over the baseline case. Increasing the duty cycle to 50% with F=0.14 and B=1.0 

appears to eliminate the separation and increases the lift to its maximum value. With higher 

frequencies a 10% duty cycle is sufficient for the same flow control, and with F=0.56 and 

D=10%, a blowing ratio of B=0.5 is sufficient. Flow control appears to be more complete and is 

possible at lower dimensionless frequencies than in the Re=25,000 cases. 

Total pressure loss results are shown in Figs. 14. Without flow control the losses are high 

and equal to those in the Re=25,000 case. With steady VGJs and B=2.0, the separation is 

NASA/CR—2012-217415 217
controlled, as shown in Fig. 13, and the loss is reduced by about 40%. It is still higher, however, 

than in some of the pulsed jet cases. With F=0.14 and B=1.0, pulsing with D=10% was shown in 

Fig. 13 to have only a limited effect controlling separation, and although the integrated total 

pressure loss is about 20% lower than in the baseline case, it is still much higher than in the other 

pulsed jet cases. The loss peaks in this case are also shifted to the left of the other VGJ cases in 

Fig. 14, indicating a reduction in flow turning. Increasing the duty cycle to 50% greatly reduces 

the losses to about 50% of the baseline case level and increases flow turning. With F=0.28, 

D=10% and B=1.0, the magnitude of the loss peaks are reduced about 40% below the baseline 

level, but the peaks are wider than in some of the other cases. Increasing F to 0.56 with B=1.0, 

the losses are reduced by 60% below the baseline level and the VGJs are equally effective with 

D=10% or 50%. The jets are slightly less effective at this frequency when B is reduced to 0.5. 

Velocity profiles for the baseline case are shown in Fig. 15. There is a large separation, and 

the dimensionless profiles are nearly identical to those in the uncontrolled case at Re=25,000. 

With steady jets at B=2.0, the profiles in Fig. 16 show that the separation bubble is eliminated 

and the boundary layer is fully turbulent at all stations. 

Time averaged velocity profiles for the case with pulsed jets at F=0.07, D=50% and B=1.0 

are shown in Fig. 17. The boundary layer remains separated, but the separation bubble is thinner 

than in the baseline case. Phase averaged results for this case are shown in Fig. 18. The 

boundary layer has separated by Station 2 and the separation bubble becomes thicker at the 

downstream stations. By Station 4 at t/T=0.167 it appears the boundary layer is close to 

reattaching. The same is true at Stations 5 at t/T=0.208 and t/T=0.667. By Station 6 the 

boundary layer appears close to reattachment at t/T=0.25 and t/T=0.708. At all other phases the 

boundary layer is clearly separated. The appearance of two times for reattachment spaced about 

a half cycle apart agrees with the results of the Re=25,000 case with D=50%, indicating that 

turning off the VGJs causes a disturbance which helps to control the boundary layer. Since no 

sign of reattachment occurs until Station 4 and the reattachment at a second phase only appears 

at Station 5, the disturbance caused by the VGJs must be too small to have an immediate effect. 

Perhaps it must grow in the separated shear layer or spread in the spanwise direction before its 

effect is visible. Although the boundary layer separates and remains separated at most phases, 

by causing the boundary layer to approach reattachment at least intermittently, the VGJs 

significantly reduce the thickness of the separation bubble and the total pressure losses. 

Velocity results with F=0.14, D=10% and B=1.0 are shown in Figs. 19 and 20. The time 

averaged results in Fig. 19 are very similar to the previous case with F=0.07, showing that the 

boundary layer separates and does not reattach. As indicated by the Cp and loss profiles of Figs. 

13 and 14, the VGJs have a relatively small effect in this case. The phase averaged results in 

Fig. 20 show that the boundary layer is near reattachment at t/T=0.25, 0.333, 0.417, and 0.5 at 

Stations 3-6 respectively. At other phases it is clearly separated. As in the previous case, the 

VGJs cause the boundary layer to approach reattachment with sufficient frequency to reduce the 

overall thickness of the separation bubble and reduce losses somewhat, but the boundary layer 

still remains separated between jet pulses. 

When the duty cycle is increased to D=50% with F=0.14 and B=1.0, the boundary layer 

reattaches, as shown in the Cp profiles and loss profiles and the time averaged velocity in Fig. 

21. The phase averaged velocity profiles in Fig. 22 show that separation is delayed until Station 

3. At phases near t/T=0.25 and 0.75 the boundary layer appears to remain attached. At other 

phases the separation bubble is thinner than in the D=10% case. The boundary layer appears 

NASA/CR—2012-217415 218
attached for about 25% of the cycle at Station 4 and about half the cycle at Station 5. By Station 

6 it is attached for about 75% of the cycle, only clearly separating near t/T=0.25. 

In Figs. 23 and 24, the pulsing frequency is increased to F=0.28 with D=10% and B=1.0. 

The boundary layer is somewhat thinner than in the previous case of Figs. 21 and 22, 

corresponding to the narrower loss peaks and lower total pressure losses in Fig. 14. As in the 

previous case, the boundary layer is attached for most of the cycle by the downstream stations. 

With the frequency increased to F=0.56 with D=10% and B=1.0, the separation bubble is 

nearly eliminated, as seen in Figs. 25 and 26. A very thin bubble appears briefly at Station 3 at 

t/T=0.5 and at Station 4 at t/T=0.708, but everywhere else the boundary layer appears attached. 

It appears close to separation at a few phases at the downstream stations, but at most phases the 

boundary layer is significantly thinner than in any of the previous cases. The fluctuating velocity 

peaks are all very close to the wall. These results are consistent with the low losses for this case. 

When the duty cycle is increased to 50% with F and B remaining at 0.56 and 1.0, Figs. 27 and 28 

show there is little change from the D=10% case. The boundary layer becomes slightly thinner, 

but the losses, as shown in Fig. 14 are about the same for the D=10% and 50% cases. Once 

F=0.56, the separation is nearly fully controlled, so increasing the duty cycle does not provide 

further benefit. Similarly, increasing the frequency to F=1.12 with D=10% and B=1.0, as shown 

in Figs. 29 and 30 causes little change. 

Figures 31 and 32 show the effect of lowering the blowing ratio to B=0.5 with F=0.56 and 

D=10%. The boundary layer separates and does not reattach. Although the separation bubble is 

much thinner and the pressure loss is much lower than in the baseline case, stronger blowing 

helps to further reduce the separation. 

Re=100,000. The Cp results for cases with Re=100,000 are shown in Figs. 33. Without flow 

control, the boundary layer separates and does not reattach, as at the lower Reynolds numbers. 

At Re=100,000, however, the boundary layer is close to a Reynolds number where it would 

reattach even without flow control. With any blowing from the VGJs, whether steady or pulsed, 

separation is controlled. Fig. 33 shows two examples, with steady blowing and B=0.75, and 

pulsed blowing with F=0.28, D=10% and B=0.5. Lift increases about 30% when the boundary 

layer reattaches. Total pressure losses are reduced by over 50% with the VGJs, as shown in Figs. 

34. 

Velocity profiles for the case without VGJs are shown in Fig. 35. The profiles match those 

without flow control at the lower Reynolds numbers. With flow control the boundary layer is 

fully attached, as shown in Fig. 36 for the case with steady jets and B=0.75 and in Figs. 37 and 

38 for F=0.28, D=10% and B=0.5. The phase averaged results show the boundary layer is 

attached at all times during the pulsing cycle. 

Effect of VGJ Spacing 

The pitch to diameter ratio of the VGJ holes was 10.7. To investigate the effect of hole 

spacing, some holes were covered with thin tape. First, before any holes were covered, velocity 

profiles were acquired at five spanwise locations at the most downstream station on the suction 

surface, s/Ls=0.97, with Re=50,000, and steady VGJs with B=2.0. These are shown in Fig. 39, 

where z is the spanwise position, and Lp is the hole pitch. The spanwise positions z/Lp=0 and 1 

correspond to the same streamwise location with respect to adjacent holes. The mean velocity 

profiles in Fig. 39a show an attached boundary layer with no variability across the span, and the 

fluctuating velocity in Fig. 39b also shows little change across the span. Hence the holes are 

NASA/CR—2012-217415 219
close enough together to provide spanwise uniform flow control. Figure 40 shows pulsed jet 

results with F=0.56, D=10% and B=1.0. Again separation is controlled and there is no variation 

across the span. The phase averaged velocity profiles (not shown) are the same as shown in Fig. 

26 for this case. A small separation does form for about 10% of the pulsing cycle, and it is 

uniform at all spanwise locations. 

Three additional hole pitches were considered. Every other hole was covered to increase the 

pitch, Lp, to Lp/d=21.3, where d is the 0.8 mm VGJ hole diameter. Every other remaining hole 

was then covered to increase the pitch to Lp/d=42.7. Finally, every other remaining hole was 

again covered to increase the pitch to Lp/d=85.3. Figure 41 shows the Cp results for these cases. 

With Lp/d=10.7 and 21.3, steady blowing with B=2.0 or pulsed blowing with F=0.56, D=10% 

and B=1.0 both cause the boundary layer to reattach. The plateau at s/Ls=0.65 indicates a larger 

separation bubble in the Lp/d=21.3 cases, but the boundary layer still reattaches well before the 

trailing edge. With the pitch increased to Lp/d=42.7, steady blowing with B=2.0 or pulsed jets 

with F=0.56, D=10% and B=1.0 are insufficient. The Cp profiles in these two cases show some 

improvement over the baseline case, but the boundary layer does not appear to reattach. With 

F=0.56, D=10% and B increased to 1.5, however, the boundary layer does reattach. With 

Lp/d=85.3, it was not possible to force reattachment. With F=0.56 and D=10%, B must be 

increased to about 2.3 to have any effect, and as shown in Fig. 41, it does not provide complete 

separation control. Raising B even more did not appear to provide further benefit. 

The total pressure loss results support the Cp measurements. Figure 42 shows results with 

Lp/d=21.3, steady jets and B=2.0. Profiles at five spanwise positions are shown. There is no 

variability across the span, and the profiles essentially match those with Lp/d=10.7 documented 

previously. The same is true for the F=0.56, D=10%, B=1.0 case shown in Fig. 43. With Lp/d 

increased to 42.7, as shown in Fig. 44 for the steady B=2.0 case and Fig. 45 for the F=0.56, 

D=10%, B=1.0 case, there is still good uniformity across the span, but the losses are much higher 

than in the cases with closer hole spacing, and the peaks are shifted to the left due to the large 

separation bubble. Figure 46 compares loss profiles acquired at z/Lp=0 with F=0.56, D=10%, 

Lp/d=42.7 and 85.3, and various blowing ratios. In agreement with the Cp results of Fig. 41, 

when Lp/d=42.7, B=1.0 is insufficient to control separation and losses are high. Increasing B to 

1.5 lowers the losses to match those in the cases with closer VGJ holes shown in Figs. 42 and 43. 

Losses are high in the case with Lp/d=85.3 even with B=2.3. 

Velocity profile measurements at s/Ls=0.97 are consistent with the pressure measurements. 

Figure 47 shows velocity profiles with Lp/d=21.3, steady jets and B=2.0. Some variation across 

the span is visible, with the profiles at z/Lp=0.25 and 0.5 having a larger velocity deficit than at 

the other locations. The variation is small, however, and separation is controlled across the span. 

The velocity deficits in Fig. 47 are lower than those in Fig. 39, indicating that increasing the hole 

spacing resulted in a thinner boundary layer. With the same hole pitch and pulsed jets with 

F=0.56, D=10% and B=1.0, Fig. 48 again shows good uniformity and separation control across 

the span. Opposite to the steady blowing case, the velocity deficit with pulsed jets is slightly 

larger than in the corresponding case with closer holes in Fig. 40. Phase averaged results show 

that a thin separation bubble forms for about 10% of the pulsing cycle, and that it is uniform 

across the span. 

With the VGJ spacing increased to Lp/d=42.7, the velocity profiles for the steady jet, B=2.0 

case are shown in Fig. 49. In agreement with the pressure results for this case, the boundary 

layer is separated. There is more variation across the span than in the cases with closer VGJs, 

but the variation is not great. The separation bubble appears thickest at z/Lp=0.25 and 0.5, and 

NASA/CR—2012-217415 220
thinnest at z/Lp=0.75. With pulsed jets at F=0.56, D=10% and B=1.0 with Lp/d=42.7, Fig. 50 

shows the boundary layer is clearly separated, again in agreement with the pressure results. In 

this case very good spanwise uniformity is maintained in spite of the VGJ spacing. Phase 

averaged results (not presented) show only a small variation in the separation bubble thickness 

during the pulsing cycle with the boundary layer remaining separated at all phases. Spanwise 

uniformity is maintained throughout the cycle. Also shown in Fig. 50 is the velocity profile from 

z/Lp=0 for the case with F=0.56, D=10%, and B=1.5. As indicated by the pressure results, 

increasing B to 1.5 causes the boundary layer to reattach. Phase averaged results show a small 

separation bubble forms for about 10% of the pulsing cycle. 

As shown above in the pressure results, with the VGJ spacing increased to Lp/d=85.3, it was 

not possible to control separation. Figure 51 shows velocity profiles from the case with F=0.56, 

D=10% and B=2.3. The boundary layer is clearly separated and there is some variation across 

the span. Phase averaged results show there is little variation with time during the pulsing cycle. 

In summary, it appears the original VGJ spacing of Lp/d=10.7 was conservative, at least for 

the Re=50,000, low freestream turbulence cases investigated on the present airfoil. The spacing 

can be doubled with little change in the flow control. If the spacing is increased to four times the 

original spacing, separation control is still possible with an increase in blowing ratio from 1.0 to 

1.5 in the pulsed jet case. There is an upper limit to the jet spacing, however, and if Lp/d is 

increased to 85.3, separation control becomes impossible. In all cases, whether separation 

control was achieved or not, good spanwise uniformity was observed near the trailing edge even 

with large VGJ spacing. This may indicate that it is necessary for the VGJs to be close enough 

together to maintain an attached boundary layer across the entire span. If the VGJs are spaced 

too far apart, strips of attached flow between regions of separation will not occur. Instead the 

boundary layer will separate across the entire span. 

VGJ Cases, High Freestream Turbulence 

Re=25,000. As documented in Volino et al. (2008), a passive grid is used to produce 

elevated freestream turbulence. At the inlet plane of the cascade, the intensity of the streamwise 

component freestream turbulence is about 4%. High freestream turbulence makes the boundary 

layer more resistant to separation, but for the present airfoil without flow control, separation 

without reattachment still occurs at Re=50,000 and below. Figure 52 shows Cp results for cases 

with Re=25,000 and steady VGJs. As in the low freestream turbulence cases at this Reynolds 

number, blowing with B=2.0 or lower has no effect on separation. When B is increased to 2.5 or 

3.0, there are some signs of reattachment, but the Cp profile remains significantly different from 

the inviscid profile and the lift remains low. 

Figures 53 shows Cp results for cases with pulsed jets and F=0.14. As in the low freestream 

turbulence cases, this frequency is too low to be effective. The jets have some effect when 

B=3.0 and the duty cycle is high, but even then the lift remains low. Figure 54 shows 

corresponding total pressure loss results. The jets do have an effect, reducing the loss 

significantly with pulsed VGJs and shifting the loss peaks to the right. This indicates that the 

separation bubble must be thinner, particularly when B=3.0, and flow turning is increased. 

The Cp results with F=0.28 are shown in Fig. 55. The jets begin to have an effect when 

B=1.5, and significant separation control is apparent when B=2.0 or higher and D=50%. The lift 

increases about 20% over the baseline case. The jets are not as effective when D=10%. The loss 

results in Fig. 56 show the VGJs at F=0.28 have a strong effect. Losses are reduced by as much 

NASA/CR—2012-217415 221
as 50%, and flow turning is increased as B increases. In contrast to the Cp results, there is little 

variation with duty cycle. Velocity profiles for the baseline case are shown in Fig. 57. The 

boundary layer has separated by Station 1 and the separation bubble grows at the downstream 

stations. Figure 58 shows time averaged profiles for the F=0.28, D=10%, B=2.0 case, and Fig. 

59 shows the corresponding phase averaged results. Boundary layer separation is delayed until 

about Station 3. At Station 3 separation is suppressed for about 25% of the cycle centered on 

t/T=0.375. This attached flow period is centered at t/T=0.583, 0.792 and 0.958 at Stations 4-6 

respectively. As noted in the low freestream turbulence cases, the fluctuating velocity peaks are 

lower during the attached flow periods. Between these periods the boundary layer separates, but 

the separation bubble is much thinner than in the baseline case. The results are essentially the 

same when B=3.0, as shown in Figs. 60 and 61. 

Figure 62 shows Cp results with F=0.56. The VGJs have some effect even at low blowing 

ratios, and when B=1.0 the boundary layer appears to reattach. The results do not appear to 

depend significantly on the duty cycle. This is consistent with the low freestream turbulence 

results, which showed that when F=0.56 the pulses occur frequently enough to control separation 

even at low duty cycles. Figure 63 shows Cp results with F=1.12. Results are similar to those 

with F=0.56. Separation is controlled when B=0.75, and increasing B further does not appear to 

have much effect. Loss results are shown in Figs. 64. Losses are greatly reduced in some of the 

cases with flow control, consistent with all the results above. There are some anomalies, 

however, comparing the loss results to the Cp profiles. A few of the cases showing reattached 

boundary layers in Figs. 62 and 64 have losses equal to the uncontrolled case in Fig. 64. 

Velocity profiles for the F=0.56, D=10%, B=1.0 case are shown in Figs. 65 and 66. The 

boundary layer separates by Station 3, except for a period near t/T=0.708. The period of 

attached flow is centered at t/T=0.958 and 0.208 at Stations 4 and 5. By Station 6 the boundary 

layer appears to be attached for most of the cycle, although the mean profile is inflectional and 

may be close to separation. The boundary layer is more attached than in the F=0.28 cases. 

Increasing B to 2.0, as shown in Figs. 67 and 68 does not significantly change the velocity 

profiles. 

Re=50,000. The Cp results with steady VGJs and Re=50,000 are shown in Fig. 69. With 

B=1.0 the VGJs appear ineffective, but with B=1.5 and 2.0 the boundary layer reattaches and the 

separation bubble appears to be eliminated. The lift increases about 16% with flow control. 

Loss results are shown in Fig. 70. In agreement with Cp, the loss peaks with B=1.0 are not much 

different than in the baseline case, but they are much lower with B=1.5 and about 40% lower 

than the baseline case with B=2.0. As the losses decrease the peaks shift to the right in Fig. 70, 

indicating more flow turning. This is consistent with the higher lift in Fig. 69b. 

Figure 71 shows Cp results for pulsed VGJs with F=0.14. As in the steady jet cases, 

separation control is best when B=1.5 or higher, and the jets are ineffective when B=0.5 or 

lower. When B=1.0, there appears to be partial separation control which is better than in the 

steady jet cases. In general the duty cycle does not affect the results, except at B=0.75 where the 

lift is higher with D=50% than with D=10%. The loss results in Fig. 72 again show some 

effectiveness with B=1.0, but the separation is not fully controlled on all blades. Results are 

better with B=1.5, with losses about half those in the baseline case at both high and low duty 

cycles. 

The Cp results with F=0.28 are shown in Fig. 73. The VGJs have some positive effect when 

B=0.5 and become more effective as B increases to 1.0. Increasing B further has little effect. 

NASA/CR—2012-217415 222
With B=0.5 and 0.75, results are better with D=50% than 10%, but at higher B the jets are 

equally effective at both duty cycles. The loss results in Fig. 74 are consistent with Cp. Velocity 

profiles for the baseline case are shown in Fig. 75 and for the F=0.28, D=10%, B=1.0 case in 

Figs. 76 and 77. The baseline results at Re=50,000 agree closely with those at Re=25,000. The 

boundary layer separates and does not reattach. With the VGJs a small separation bubble may 

form between Station 3 and 4, and the boundary layer appears to be approaching separation at 

the downstream stations at some phases (e.g. t/T=0.583 at Station 6), but for the most part the 

boundary layer remains attached. 

With F=0.56, Fig. 78 shows the Cp results. Separation is controlled with B=0.75 or higher, 

independent of duty cycle. Higher duty cycle results in higher lift when B=0.5. Loss results in 

Fig. 79 are consistent with the Cp results. Velocity profiles in Figs. 80 and 81 for the F=0.56, 

D=10%, B=0.75 case show that the boundary layer remains attached for the full pulsing cycle. 

Summary of VGJ Cases 

Boundary layer separation was controlled using VGJs at all Reynolds numbers considered. 

The disturbance caused when the jets are turned on or off is more effective for flow control than 

steady blowing. With a low duty cycle of 10%, the switching on and off of the jets occur close 

enough together in time that they provided essentially a single disturbance each pulsing cycle. 

Hence the disturbance frequency is the same as the pulsing frequency. When the duty cycle is 

increased to 50%, the switching on and off are separated by half a cycle and provide two 

disturbances per cycle. This effectively doubles the disturbance frequency over the pulsing 

frequency. A strong enough disturbance from the VGJs causes the boundary layer to either 

remain attached or reattach farther downstream. After the disturbance passes, the boundary layer 

can relax and eventually re-separate. As the disturbance frequency is increased, the time 

available for separation between disturbances shrinks, allowing the separation bubble less time to 

grow. This reduces the average size of the bubble, resulting in greater flow turning, higher lift, 

and lower losses. If the disturbance frequency is sufficiently high, the separation bubble can be 

eliminated in some cases. The acceleration parameter, K=(ν/U∞

2

)(dU∞/ds), is an indicator of the 

tendency of the boundary layer to separate, and is inversely proportional to the Reynolds 

number. At lower Reynolds numbers, therefore, the boundary layer should require a smaller 

fraction of a disturbance cycle to separate, and a higher dimensionless pulsing frequency may be 

needed to control separation. 

The velocity of the VGJs must also be sufficiently high to provide separation control. Since 

steady jets are less effective, they require a higher blowing ratio to produce the same effect as a 

pulsed jet. With a higher blowing ratio and 100% duty cycle, the steady jets use considerably 

more air than the pulsed jets, which produces higher overall losses. With pulsed jets, the 

necessary blowing ratio depends on the pulsing frequency. Presumably this is because at lower 

frequencies a strong pulse is needed to more fully reattach the boundary layer and fill out the 

velocity profile near the wall, thereby lengthening the time needed after the disturbance for 

separation. With higher frequencies the boundary layer has less time to separate, so a weaker 

pulse may be sufficient. If the VGJs are not strong or frequent enough to prevent separation or 

force the boundary layer to reattach, they may still be able to reduce the separation bubble 

thickness. This is particularly true at low Reynolds numbers, where the boundary layer is very 

prone to separation and remains thick and on the verge of separation at the downstream stations 

even in the best cases. 

NASA/CR—2012-217415 223
The discussion above is illustrated in terms of the integrated total pressure loss in Fig. 82. 

The loss is shown as a function of F and B. Data were not acquired at every possible 

combination of F and B, so considerable interpolation was used to fill in the contour plots. 

Quantitative information should therefore be taken from the figures discussed above instead of 

Fig. 82. Figure 82a shows that with low freestream turbulence and Re=25,000, a pulsing 

frequency of about F=0.5 is needed, and a blowing ratio of about B=1.0 is sufficient if the 

frequency is high enough. Further increasing F or B provides some additional benefit. With 

lower frequencies the VGJs are not helpful in terms of losses. The high blowing ratios needed to 

provide any separation control result in higher losses from the jets themselves than the benefit 

gained from reducing the separation bubble thickness. 

With low freestream turbulence and Re=50,000, Fig. 82b shows that with F=0.5, low losses 

are possible with B as low as 0.25. With F=0.3, similar losses are possible with B of about 1.0. 

The separation is controlled to a sufficient extent in these cases that additional increases in F or 

B do not provide further benefit. At lower frequencies the separation control is not as good and 

losses are higher even at the highest blowing rates considered. 

The addition of high freestream turbulence does not drastically change the results. When 

Re=100,000, high freestream turbulence causes the boundary layer to reattach without the use of 

flow control. At low Reynolds numbers flow control is still needed to induce reattachment. 

When Re=25,000, as shown in Fig. 82c, a frequency of about F=0.3 and blowing ratio of about 

B=1.0 are needed to significantly reduce losses, and higher B is helpful. When Re=50,000, as 

shown in Fig. 82d, a frequency of F=0.3 and blowing ratio of about 0.75 are sufficient, and 

higher frequency and blowing ratio do not appear to reduce losses further. 

CFD for VGJ Cases 

Previous work (Ibrahim et al., 2008, Volino et al., 2008) showed that the URANS 

calculations using the Transition-sst model were successful in predicting separation, transition 

and reattachment accurately in the baseline cases without flow control. Several other models 

were also tested. Many failed to predict separation. Some others did better, but not as well as 

the Transition-sst model. Pressure distributions and velocity profiles from the CFD and 

experiments were compared. The calculations matched the experiments, and correctly predicted 

the effects of freestream turbulence at both high and low Reynolds numbers. Also shown in 

Volino et al. (2009) were attempts to apply the same URANS methods with the Transition-sst 

model to cases with VGJ flow control. The calculations failed to predict the effect of the VGJs. 

Reattachment was not predicted, even in cases where the VGJs almost completely eliminated the 

separation in the experiments. To investigate the problem and see whether a good CFD 

prediction could be obtained by another method, large eddy simulations (LES) were done for 

some cases. The results are documented in Ibrahim et al. (2009) which is included at the end of 

this report. The match between the LES and experiments was not perfect, but LES was able to 

predict reattachment in the correct cases. It appears the much finer grid required for the LES was 

needed to capture the effect of the VGJs. The disadvantage of LES is the long computing time 

required due to the very fine grid. Further work is planned to determine how fine a grid is 

needed for accurate predictions with URANS. 

NASA/CR—2012-217415 224
Wakes Cases, Low Freestream Turbulence, No VGJs 

Wake Characteristics. Cases without VGJ flow control were documented at Re=25,000 

and 50,000 using the wake generator described above. The wakes generated by the moving rods 

can be characterized by comparing them to the wakes from the blades in the cascade. Figure 83 

shows the mean and fluctuating velocity in the wakes of the moving rods and of the cascade 

blades. The cascade wakes were measured by traversing a hot-wire probe in lines 0.3Cx and 

0.63Cx downstream of the trailing edges of the blades, perpendicular to the axial direction. In 

the 0.63Cx case, this is the same location used for the total pressure loss measurements. The 

spacing between airfoil rows in an engine is expected to be about 0.3Cx according to Pluim et al. 

(2009), so this location would be where the wake from an upstream airfoil would reach the 

leading edge of the next row. The moving rod wakes were measured by placing a stationary hotwire midway between two blades in the cascade in the plane of the leading edges. The moving 

rods traversed 0.54Cx upstream of the hot-wire. The data from the probe were phase averaged on 

the rod passing cycle to determine the mean and fluctuating velocity as functions of time. For 

comparison in the figures, the rod velocity was then used to convert time to distance traversed. 

Curves in the figures were shifted along the horizontal axis to align the peaks for comparison. 

The actual positions, as shown previously, depend on the actual amount of flow turning, which 

varies with boundary layer separation and is a function of Reynolds number. All curves were 

normalized using the average local mean velocity. 

The wakes from the upstream rods, when shown in dimensionless form as in Fig. 83, did not 

depend on Reynolds number. For all of the flow and rod velocities considered, the Reynolds 

number based on rod diameter is between about 400 and 2000, which should result in laminar 

separation for the boundary layer on the rod. With all cases in the same flow regime, the wakes 

should be similar. The wake for the Re=50,000 case with dimensionless wake passing frequency 

of F=0.28 is shown. For consistence, F is defined as above for the pulsed jet cases using the 

distance from the VGJ holes to the trailing edge and the average freestream velocity over this 

distance. In contrast to the rod wakes, the wakes of the cascade airfoils depend strongly on 

Reynolds number. As shown in Volino (2008) there are two flow regimes for the wakes. For 

Re=100,000 and below, the boundary layer separates and does not reattach. The wake is wide, 

the velocity deficit is large, and the peak turbulence intensity is of the order 20%. In 

dimensionless form, all cases in this regime are quantitatively similar. The second regime 

corresponds to Re=200,000 and above. The wakes in this regime are much narrower with lower 

velocity deficits and peak turbulence intensities of about 5%. In dimensionless form, all cases in 

this regime are also quantitatively similar. Figure 83 shows one case from each regime at 

Re=50,000 and 200,000. 

The airfoil wake at Re=50,000 has a large mean velocity deficit at 0.3Cx which decreases in 

peak magnitude by 80% by 0.63Cx. The peak turbulence intensity drops by 40% over this 

distance. At Re=200,000, the airfoil wake is about half as wide as at the lower Re, and the peak 

velocity deficit is about 25% as large. The peak turbulence intensity is 7%, which is about a 

third that at the lower Re. The change in the wake between 0.3Cx and 0.63Cx is much smaller 

than at the lower Re. The rod wake lies between the high and low Re airfoil wakes. The peak 

turbulence intensity is 14%. The rod wake appears to be reasonable approximation to an airfoil 

wake, although it does not fully match either the high or low Re case. One could conceivably 

use a rod of either larger or smaller diameter to match the low or high Re airfoil cases better. 

NASA/CR—2012-217415 225
Also of interest is how the rod wake compares to the airfoil wake in a low Reynolds number 

case with VGJs used to successfully control separation. The Re=50,000, F=0.56, B=1.0 case is 

used for comparison. As shown in Fig. 26, the boundary layer separates very briefly in this case 

at the upstream stations, but reattaches. At the most downstream station the boundary layer is 

attached for the full pulsing cycle, although it appears to be tending toward re-separation for part 

of the cycle. Figure 84 shows the wake velocity for this case and compares it to the rod wake. 

For the airfoil wake, the phase averaged results at 24 phases through the pulsing cycle are shown. 

Since the boundary layer does approach separation during part of the cycle, the wake velocity 

profile magnitude and position changes somewhat during the cycle. The mean velocity deficit 

and wake width of the upstream rod and airfoil are about the same. The peak turbulence 

intensity in the rod wake is about 40% higher than the highest value in the airfoil wake. 

Turbulence spectra provide additional information about the wakes. The Fourier spectra of 

the airfoil wakes for the baseline cases are shown in Fig. 85. Dimensionless frequency on a log 

scale and position across the cascade at 0.63Cx downstream of the trailing edges are on the 

horizontal axes. The dimensionless frequency, F, is defined as above for the pulsed jet cases. 

The vertical axis shows the magnitude of the power spectral density premultiplied by the 

frequency and normalized on the nominal cascade exit velocity squared. In these coordinates the 

area under the surface at any position and frequency is proportional to the energy in the 

turbulence at that position and frequency. For the Re=50,000 case in Fig. 85a, there are peaks 

for each wake located at F=0.7 and smaller peaks at F=1.4. The expected vortex shedding 

frequency for the separation bubble on the airfoils, based on a bubble thickness of about 0.02 m, 

the trailing edge flow velocity, and a Strouhal number of 0.2 is about F=1.0. The observed 

frequencies in the wake are therefore consistent with vortex shedding. The spectra for the 

Re=200,000 case in Fig. 85b are much different. The peaks are two orders of magnitude lower 

than in the Re=50,000 case. This difference in magnitude of u′

2

 is consistent with the one order 

of magnitude difference in u′ shown in Fig. 83b. The spectral peaks are spatially narrower at 

Re=200,000, again consistent with the wake velocity profiles in Fig. 83. The spectral peaks are 

broad banded and centered at F=2.5. These peaks are likely associated with the turbulence in the 

airfoil boundary layers. Similar peaks at F=2.5 may also be present in the Re=50,000 case, but if 

they are present they are overwhelmed in the tails of the much larger low frequency peaks. Low 

frequency peaks are not observed at Re=200,000 because the airfoil boundary layers are attached 

so there is no large separation bubble to produce vortex shedding. 

Fourier spectra for the cascade wakes with pulsed VGJs at Re=50,000, F=0.56, D=10%, and 

B=1.0 are shown in Fig. 86. The spectra have magnitude and frequency between those of the 

high and low Re cases of Fig. 85. Fourier spectra for this case present an incomplete picture, 

however, because the wake moves spatially during the jet pulsing cycle and the magnitude of the 

peaks also change. Fourier spectra present a time average of the quantity under consideration, so 

it is not possible to consider frequencies at particular times during the cycle. To overcome this, 

wavelet spectra can be computed. In wavelet analysis, the convolution integral between a signal 

of interest and a function containing oscillations at a particular frequency (the wavelet) is 

computed. The magnitude of the result indicates when in time the frequency of the wavelet 

appears in the signal. The wavelet is dilated and constricted to change its oscillation frequency 

and the convolution integral is computed multiple times to construct a map of the signal content 

as a function of time and frequency simultaneously. The particular wavelet, or function, used for 

the analysis determines how well time and frequency are each resolved. Two commonly used 

wavelets for turbulence analysis are the Mexican Hat wavelet and the Morlet wavelet. The 

NASA/CR—2012-217415 226
Mexican Hat wavelet is a real valued function with essentially a single oscillation along the time 

axis. It is particularly good for resolving oscillations in time, but less so for frequency. The 

Morlet wavelet is a complex valued with multiple oscillations at a given frequency along the 

time axis. Because it has a longer duration, it is less able to resolve time, but because of its 

multiple oscillations it can better resolve frequencies than the Mexican Hat wavelet. Figure 87 

shows the wavelet spectra for the wake turbulence of the pulsed jet case at two representative 

times in the cycle. Figures 87a and 87b were produced with the Mexican Hat wavelet and Figs. 

87c and 87d show the same data analyzed with the Morlet wavelet. Figure 87a is typical of most 

of the cycle and shows a peak for each wake centered at F=1.9. There is also a smaller peak, 

hidden in the view shown at F=0.56 at a pitchwise position about 0.4Lφ higher than the large 

peak. The smaller peak is at the pulsing frequency. At the phase shown in Fig. 87b, the higher 

frequency peak magnitude is now about a third of its value in Fig. 87a and is split into two 

nearby peaks at F=1.6 and 2.2. The F=0.56 peak is still present at about the same magnitude, but 

now appears more prominent since the other peaks are smaller. If all phases of the wavelet 

spectra are time averaged, the result is close to the Fourier spectra of Fig. 86. Figure 87c shows 

the Morlet wavelet result at the same phase shown in Fig. 87a. The two wavelets produce 

similar results, but the low frequency peak at the pulsing frequency is better resolved and has a 

higher magnitude with the Morlet wavelet. Figure 87d shows the Morlet wavelet result at the 

same phase as Fig. 87b. Again the two wavelets produce similar results, with some differences 

in the peak magnitude and the Morlet wavelet showing a sharper peak at the pulsing frequency. 

Figure 88a shows the rod wake spectrum using the Morlet wavelet. The time axis is 

converted to a spatial position, as explained above for Fig. 83. There is a sharp peak at F=0.28, 

which corresponds to the rod passing frequency. A broader peak is centered at about F=1.0. 

This broad peak is due to the turbulence and vortices shed by the rod. Figure 88b shows the 

same data analyzed with the Mexican Hat wavelet. No peak appears at the wake passing 

frequency because the Mexican Hat wavelet is more narrowly focused in time and does not pick 

up the variation in velocity across the cycle. The broad band peak is centered at F=0.8, which is 

slightly different than indicated by the Morlet wavelet. It is also somewhat narrower spatially 

due to the better time resolution of the Mexican Hat wavelet. As with the mean and rms 

velocity, the spectra of the rod wakes have frequency and magnitude between those of the high 

and low Reynolds number baseline cases. The rod wakes are more similar to those of the pulsed 

jet controlled case of Fig. 87. The peak magnitudes are similar, although the frequencies of the 

rod wakes are somewhat lower. Given that there are a variety of wakes that could be generated 

by an upstream airfoil depending on the flow conditions, the wakes produced by the present 

cylindrical rods appear to be reasonable for the present investigation. Investigating the effect of 

different wakes produced by rods of different diameter or shape may be an interesting topic for 

future work. 

Pressure, loss, and velocity results. The Cp distributions for the cases with wakes are 

shown in Fig. 89. With the wake passing period set to F=0.28, the wakes appear to have some 

effect at Re=25,000, but the boundary layer does not appear to reattach. At Re=37,500 the Cp 

values are closer to the inviscid line near the trailing edge, suggesting possible reattachment. In 

the Re=50,000 case there is good agreement with the inviscid line, suggesting the boundary layer 

is reattached. If the wake passing frequency is doubled to F=0.56 in the Re=25,000 case, the Cp 

values indicate more reattachment, although the suction peak is still low compared to the higher 

Re cases. When the wake passing frequency is lowered to F=0.14 with Re=50,000, the wakes 

NASA/CR—2012-217415 227
have some effect in lowering Cp, but the boundary layer appears to remain separated. The 

boundary layer reattaches without flow control when Re=100,000 in the high freestream 

turbulence case and at Re slightly above 100,000 with low freestream turbulence. It is not 

surprising, therefore, that with Re=100,000 there is good agreement with the inviscid line even 

with the low wake passing frequency of F=0.14. When Re=200,000 there is no separation 

problem in the baseline cases, so adding wakes, regardless of the frequency does not change the 

Cp distribution. The wake passing frequency needed to suppress separation appears to match the 

pulsing frequency needed for control with VGJs at moderate blowing ratios. At Re=50,000, 

F=0.28 is sufficient, but at lower Reynolds numbers a higher frequency is needed. 

Figure 90 shows the total pressure loss results. With F=0.28, the wakes have an effect at all 

Reynolds numbers, suppressing the loss peak and establishing periodicity across the cascade. At 

Re=25,000 the peak is wide, indicating a thick boundary layer. Results were similar with VGJs 

at this frequency. As Re is increased, the peak becomes narrower, indicating a thinner boundary 

layer and lower overall losses. When F is increased to 0.56, the peak becomes somewhat 

narrower at Re=25,000, but it is still wider than in the higher Re cases. Lowering F to 0.14 

causes a large increase in losses and a loss of periodicity in the cascade when Re=50,000. This is 

consistent with the Cp results. At Re=100,000 the wakes suppress separation much like the 

VGJs, and the loss peaks match those obtained with VGJs. At Re=200,000 the loss peaks match 

those in the baseline case. Adding wakes or changing their frequency does not appear to have an 

effect at high Reynolds number. Any losses induced by the wakes themselves must be relatively 

small. 

Figure 91 shows velocity profiles for the Re=25,000, F=0.28 case. Data were not acquired at 

Station 1. At Station 2, a disturbance in the mean profiles appears between t/T=0.375 and 0.583 

due to the wake. By Station 3 the boundary layer has separated, although the separation appears 

to be suppressed between t/T=0.458 and 0.583. The separation bubble grows bigger at some 

phases at Stations 4 and 5, but it is still suppressed at some phases. At Station 5 the wake 

appears to be reducing the extent of the separation for about half the cycle. At Station 6 the 

boundary layer has an inflection point at all phases and appears on the verge of separating, but is 

only clearly separated for a few phases centered around t/T=0.542. The wakes significantly 

suppress the separation bubble compared to the baseline case, but do not eliminate it entirely. 

This is similar to the effect of VGJs at this Reynolds number. The result is losses much lower 

than the baseline case, but still significantly higher than with higher Reynolds numbers due to the 

thick boundary layer. 

Figure 92 shows the results with Re=37,500 and F=0.28. In agreement with the Cp and loss 

profiles, results are similar to those at Re=25,000, but the separation bubble thickness is reduced. 

With Re=50,000 and F=0.28, as shown in Fig. 93, the trend continues to the point that the 

separation bubble is essentially eliminated. With F=0.14 at Re=50,000 in Fig. 94, the separation 

bubble reappears. The wake passing frequency is not high enough to keep the boundary layer 

attached. 

Timeline and Upcoming Work 

Work was delayed in the second half of 2008 so that the wind tunnel could be disassembled 

for fabrication and installation of the wake generator. This was completed in November 2008 

and velocity data with VGJs and low freestream turbulence was acquired in late 2008 and early 

2009. The turbulence grid was then installed, and pressure and velocity data were acquired for 

NASA/CR—2012-217415 228
high freestream turbulence cases with VGJs. Experiments with the wake generator began in 

April, and low freestream turbulence cases with wakes were documented in May and June. 

Computations of cases with VGJs showed that URANS calculations with the Transition-sst 

model did not predict the flow control observed in the experiments. Large eddy simulations 

were done in the first part of 2009 to try to produce an accurate CFD prediction. The LES 

correctly predicted boundary layer reattachment. 

Work is currently somewhat behind the original proposed schedule due to the late start at the 

very beginning of the project which resulted from problems with transfer of funds, and necessary 

pauses in the experiments for fabrication and repair of the facility, and the need to pursue 

different approaches with the CFD. Work is, however, proceeding as planned. Experiments 

with high freestream turbulence and wakes will be done next, followed by cases with wakes and 

VGJs combined. CFD with wakes is beginning now, and efforts to simulate cases with VGJs 

using URANS calculations continue. 

References 

Ibrahim, M., Kartuzova, O., and Volino, R.J., 2008, “Experimental and Computational 

Investigations of Separation and Transition on a Highly Loaded Low-Pressure Turbine Airfoil: 

Part 1 – Low Freestream Turbulence Intensity,” ASME paper IMECE2008-68879. Presented at 

the ASME IMECE, Boston, MA, November 2008. In press for ASME J. Turbomachinery. 

Ibrahim, M., Kartuzova, O., and Volino, R., 2009, “LES and URANS Computational 

Investigations of LPT Blade (L1A) Separation Control using Vortex Generator Jets,” accepted 

for presentation at Turbulence, Heat and Mass Transfer 6, Rome, Italy, September 2009. 

Pluim, J., Memory, C., Bons, J., and Chen, J.P., “Designing a High Fidelity Wake Simulator 

for Research Using Linear Cascades,” ASME paper GT2009-59276. Presented at the ASME 

International Gas Turbine Conference, Orlando, FL, June 2009. 

Volino, R.J., 2008, “Separated Flow Measurements on a Highly Loaded Low-Pressure 

Turbine Airfoil,” ASME paper GT2008-51445. Presented at the ASME International Gas 

Turbine Conference, Berlin, Germany, June 2008. 

Volino, R.J., Kartuzova, O., and Ibrahim, M., 2008, “Experimental and Computational 

Investigations of Separation and Transition on a Highly Loaded Low-Pressure Turbine Airfoil: 

Part 2 – High Freestream Turbulence Intensity,” ASME paper IMECE2008-68776. Presented at 

the ASME IMECE, Boston, MA, November 2008. 

Volino, R.J., Kartuzova, O., and Ibrahim, M.B., 2009, “Experimental and Computational 

Investigations of Low-Pressure Turbine Separation Control Using Vortex Generator Jets,” 

ASME paper GT2009-59983. Presented at the ASME International Gas Turbine Conference, 

Orlando, FL, June 2009. 

NASA/CR—2012-217415 229
0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

0

0.5

1

1.5

2

2.5

s/L

s

Cp

No jets

Steady jets, B=3.0

F=0.28 D=10 B=1.0

F=0.28 D=50 B=1.0

F=0.56 D=10 B=1.0

F=1.12 D=10 B=1.0

Inviscid

a) 

NASA/CR—2012-217415 230
0 0.2 0.4 0.6 0.8 1 1.2 1.4

0.8

0.85

0.9

0.95

1

1.05

1.1

1.15

F

Cp/Cpinv

No jets

Steady jets, B=3.0

Pulsed jets, B=1.0

b) 

Fig. 1 Cp results, low TI, Re=25,000: a) Cp profile, b) Integrated Cp. 

NASA/CR—2012-217415 231
−1.5 −1 −0.5 0 0.5 1 1.5

0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4

B5 B6

No jets

Steady jets, B=3.0

F=0.28 D=10 B=1.0

F=0.56 D=10 B=1.0

F=1.12 D=10 B=1.0

a) 

NASA/CR—2012-217415 232
0 0.2 0.4 0.6 0.8 1 1.2 1.4

0.35

0.4

0.45

0.5

0.55

0.6

0.65

0.7

F

ψint

No jets

Steady jets, B=3.0

Pulsed jets, B=1.0

b) 

Fig. 2 Ψ results, low TI, Re=25,000: a) Ψ profile, b) Integrated Ψ. 

0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

NASA/CR—2012-217415 233
Fig. 3 Time averaged velocity profiles, low TI, Re=25,000, no VGJs: a) mean, b) rms. 

0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 4 Time averaged velocity profiles, low TI, Re=25,000, steady VGJs, B=3.0: a) mean, b) 

rms. 

0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 5 Time averaged velocity profiles, low TI, Re=25,000, pulsed VGJs, F=0.28, D=10%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 234
0 0.1 0.2

0.042

Re=25,000, f=06, d=10, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

a) 

NASA/CR—2012-217415 235
0 0.1 0.2

0.042

Re=25,000, f=06, d=10, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/Ue

b) 

Fig. 6 Phase averaged velocity profiles, low TI, Re=25,000, pulsed VGJs, F=0.28, D=10%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 236
0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 7 Time averaged velocity profiles, low TI, Re=25,000, pulsed VGJs, F=0.28, D=50%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 237
0 0.1 0.2

0.042

Re=25,000, f=06, d=50, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

a) 

NASA/CR—2012-217415 238
0 0.1 0.2

0.042

Re=25,000, f=06, d=50, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/Ue

b) 

Fig. 8 Phase averaged velocity profiles, low TI, Re=25,000, pulsed VGJs, F=0.28, D=50%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 239
0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 9 Time averaged velocity profiles, low TI, Re=25,000, pulsed VGJs, F=0.56, D=10%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 240
0 0.1 0.2

0.042

Re=25,000, f=12, d=10, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

a) 

NASA/CR—2012-217415 241
0 0.1 0.2

0.042

Re=25,000, f=12, d=10, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/Ue

b) 

Fig. 10 Phase averaged velocity profiles, low TI, Re=25,000, pulsed VGJs, F=0.56, D=10%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 242
0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 11 Time averaged velocity profiles, low TI, Re=25,000, pulsed VGJs, F=1.12, D=10%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 243
0 0.1 0.2

0.042

Re=25,000, f=24, d=10, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

a)

NASA/CR—2012-217415 244
0 0.1 0.2

0.042

Re=25,000, f=24, d=10, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/Ue

b) 

Fig. 12 Phase averaged velocity profiles, low TI, Re=25,000, pulsed VGJs, F=1.12, D=10%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 245
0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

0

0.5

1

1.5

2

2.5

s/L

s

Cp

No jets

Steady jets, B=2.0

F=0.14 D=10 B=1.0

F=0.14 D=50 B=1.0

F=0.28 D=10 B=1.0

F=0.56 D=10 B=0.5

F=0.56 D=10 B=1.0

F=0.56 D=50 B=1.0

Inviscid

a) 

NASA/CR—2012-217415 246
0 0.1 0.2 0.3 0.4 0.5 0.6 0.7

0.8

0.85

0.9

0.95

1

1.05

F

Cp/Cpinv

No jets

Steady jets, B=2.0

Pulsed jets, D=10 B=1.0

Pulsed jets, D=50 B=1.0

Pulsed jets, D=10 B=0.5

b) 

Fig. 13 Cp results, low TI, Re=50,000: a) Cp profile, b) Integrated Cp. 

NASA/CR—2012-217415 247
−1.5 −1 −0.5 0 0.5 1 1.5

0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4

B5 B6

No jets

Steady jets, B=2.0

F=0.14 D=10 B=1.0

F=0.14 D=50 B=1.0

F=0.28 D=10 B=1.0

F=0.56 D=10 B=0.5

F=0.56 D=10 B=1.0

F=0.56 D=50 B=1.0

a) 

NASA/CR—2012-217415 248
0 0.1 0.2 0.3 0.4 0.5 0.6 0.7

0.2

0.25

0.3

0.35

0.4

0.45

0.5

F

ψint

No jets

Steady jets, B=2.0

Pulsed jets, D=10 B=1.0

Pulsed jets, D=50 B=1.0

Pulsed jets, D=10 B=0.5

b) 

Fig. 14 Ψ results, low TI, Re=50,000: a) Ψ profile, b) Integrated Ψ. 

NASA/CR—2012-217415 249
0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 15 Time averaged velocity profiles, low TI, Re=50,000, no VGJs: a) mean, b) rms. 

0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 16 Time averaged velocity profiles, low TI, Re=50,000, steady VGJs, B=2.0: a) mean, b) 

rms. 

NASA/CR—2012-217415 250
0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 17 Time averaged velocity profiles, low TI, Re=50,000, pulsed VGJs, F=0.07, D=50%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 251
0 0.1 0.2

0.042

Re=50,000, f=03, d=50, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

a) 

NASA/CR—2012-217415 252
0 0.1 0.2

0.042

Re=50,000, f=03, d=50, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/Ue

b) 

Fig. 18 Phase averaged velocity profiles, low TI, Re=50,000, pulsed VGJs, F=0.07, D=50%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 253
0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 19 Time averaged velocity profiles, low TI, Re=50,000, pulsed VGJs, F=0.14, D=10%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 254
0 0.1 0.2

0.042

Re=50,000, f=06, d=10, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

a) 

NASA/CR—2012-217415 255
0 0.1 0.2

0.042

Re=50,000, f=06, d=10, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/Ue

b) 

Fig. 20 Phase averaged velocity profiles, low TI, Re=50,000, pulsed VGJs, F=0.14, D=10%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 256
0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 21 Time averaged velocity profiles, low TI, Re=50,000, pulsed VGJs, F=0.14, D=50%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 257
0 0.1 0.2

0.042

Re=50,000, f=06, d=50, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

a) 

NASA/CR—2012-217415 258
0 0.1 0.2

0.042

Re=50,000, f=06, d=50, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/Ue

b) 

Fig. 22 Phase averaged velocity profiles, low TI, Re=50,000, pulsed VGJs, F=0.14, D=50%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 259
0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 23 Time averaged velocity profiles, low TI, Re=50,000, pulsed VGJs, F=0.28, D=10%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 260
0 0.1 0.2

0.042

Re=50,000, f=12, d=10, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

a) 

NASA/CR—2012-217415 261
0 0.1 0.2

0.042

Re=50,000, f=12, d=10, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/Ue

b) 

Fig. 24 Phase averaged velocity profiles, low TI, Re=50,000, pulsed VGJs, F=0.28, D=10%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 262
0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 25 Time averaged velocity profiles, low TI, Re=50,000, pulsed VGJs, F=0.56, D=10%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 263
0 0.1 0.2

0.042

Re=50,000, f=24, d=10, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

a) 

NASA/CR—2012-217415 264
0 0.1 0.2

0.042

Re=50,000, f=24, d=10, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/Ue

b) 

Fig. 26 Phase averaged velocity profiles, low TI, Re=50,000, pulsed VGJs, F=0.56, D=10%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 265
0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 27 Time averaged velocity profiles, low TI, Re=50,000, pulsed VGJs, F=0.56, D=50%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 266
0 0.1 0.2

0.042

Re=50,000, f=24, d=50, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

a) 

NASA/CR—2012-217415 267
0 0.1 0.2

0.042

Re=50,000, f=24, d=50, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/Ue

b) 

Fig. 28 Phase averaged velocity profiles, low TI, Re=50,000, pulsed VGJs, F=0.56, D=50%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 268
0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 29 Time averaged velocity profiles, low TI, Re=50,000, pulsed VGJs, F=1.12, D=10%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 269
0 0.1 0.2

0.042

Re=50,000, f=48, d=10, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

a) 

NASA/CR—2012-217415 270
0 0.1 0.2

0.042

Re=50,000, f=48, d=10, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/Ue

b) 

Fig. 30 Phase averaged velocity profiles, low TI, Re=50,000, pulsed VGJs, F=1.12, D=10%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 271
0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 31 Time averaged velocity profiles, low TI, Re=50,000, pulsed VGJs, F=0.56, D=10%, 

B=0.5: a) mean, b) rms. 

NASA/CR—2012-217415 272
0 0.1 0.2

0.042

Re=50,000, f=24, d=10, b=0.5

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

a) 

NASA/CR—2012-217415 273
0 0.1 0.2

0.042

Re=50,000, f=24, d=10, b=0.5

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/Ue

b) 

Fig. 32 Phase averaged velocity profiles, low TI, Re=50,000, pulsed VGJs, F=0.56, D=10%, 

B=0.5: a) mean, b) rms. 

NASA/CR—2012-217415 274
0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

0

0.5

1

1.5

2

2.5

s/L

s

Cp

No jets

Steady jets, B=0.75

F=0.28 D=10 B=0.5

Inviscid

a) 

NASA/CR—2012-217415 275
0 0.05 0.1 0.15 0.2 0.25 0.3 0.35

0.75

0.8

0.85

0.9

0.95

1

1.05

1.1

1.15

F

Cp/Cpinv

No jets

Steady jets, B=0.75

Pulsed jets, D=10 B=0.5

b) 

Fig. 33 Cp results, low TI, Re=100,000: a) Cp profile, b) Integrated Cp. 

NASA/CR—2012-217415 276
−1.5 −1 −0.5 0 0.5 1 1.5

0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4

B5 B6

No jets

Steady jets, B=0.75

a) 

NASA/CR—2012-217415 277
0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8

0.2

0.25

0.3

0.35

0.4

F

ψint

No jets

Steady jets, B=0.75

b) 

Fig. 34 Ψ results, low TI, Re=100,000: a) Ψ profile, b) Integrated Ψ. 

NASA/CR—2012-217415 278
0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 35 Time averaged velocity profiles, low TI, Re=100,000, no VGJs: a) mean, b) rms. 

0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 36 Time averaged velocity profiles, low TI, Re=100,000, steady VGJs, B=0.75: a) mean, b) 

rms. 

NASA/CR—2012-217415 279
0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 37 Time averaged velocity profiles, low TI, Re=100,000, pulsed VGJs, F=0.28, D=10%, 

B=0.5: a) mean, b) rms. 

NASA/CR—2012-217415 280
0 0.1 0.2

0.042

Re=100,000, f=24, d=10, b=0.5

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

a) 

NASA/CR—2012-217415 281
0 0.1 0.2

0.042

Re=100,000, f=24, d=10, b=0.5

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/Ue

b) 

Fig. 38 Phase averaged velocity profiles, low TI, Re=100,000, pulsed VGJs, F=0.28, D=10%, 

B=0.5: a) mean, b) rms. 

NASA/CR—2012-217415 282
0 0.05 0.1 0.15 0.2 0.25

0

0.2

0.4

0.6

0.81

y/L

s

U/U

e

z/L

p

=0

z/L

p

=0.25

z/L

p

=0.5

z/L

p

=0.75

z/L

p

=1.0

a) 

NASA/CR—2012-217415 283
0 0.05 0.1 0.15 0.2 0.25

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

0.2

y/Ls

u′/Ue

z/L

p

=0

z/L

p

=0.25

z/L

p

=0.5

z/L

p

=0.75

z/L

p

=1.0

b) 

Fig. 39 Time averaged velocity profiles, low TI, Re=50,000, steady VGJs, B=2.0, Lp/d=10.7: a) 

mean, b) rms. 

NASA/CR—2012-217415 284
0 0.05 0.1 0.15 0.2 0.25

0

0.2

0.4

0.6

0.81

y/L

s

U/U

e

z/L

p

=0

z/L

p

=0.25

z/L

p

=0.5

z/L

p

=0.75

z/L

p

=1.0

a) 

NASA/CR—2012-217415 285
0 0.05 0.1 0.15 0.2 0.25

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

0.2

y/Ls

u′/Ue

z/L

p

=0

z/L

p

=0.25

z/L

p

=0.5

z/L

p

=0.75

z/L

p

=1.0

b) 

Fig. 40 Time averaged velocity profiles, low TI, Re=50,000, pulsed VGJs, F=0.56, D=10%, 

B=1.0, Lp/d=10.7: a) mean, b) rms. 

NASA/CR—2012-217415 286
0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

0

0.5

1

1.5

2

2.5

3

s/L

s

Cp

Re=50,000

No jets

L

p

/d=10.7 F=0 B=2.0

L

p

/d=10.7 F=0.56 D=10 B=1.0

L

p

/d=21.3 F=0 B=2.0

L

p

/d=21.3 F=0.56 D=10 B=1.0

L

p

/d=42.7 F=0 B=2.0

L

p

/d=42.7 F=0.56 D=10 B=1.0

L

p

/d=42.7 F=0.56 D=10 B=1.5

L

p

/d=85.3 F=0.56 D=10 B=2.3

Inviscid

a) 

NASA/CR—2012-217415 287
0 10 20 30 40 50 60 70 80 90

0.8

0.85

0.9

0.95

1

1.05

1.1

1.15

1.2

1.25

L

p

/d

Cp/Cpinv

No jets

Steady jets, B=2.0

Pulsed jets, F=0.56 D=10 B=1.0

Pulsed jets, F=0.56 D=10 B=1.5

Pulsed jets, F=0.56 D=10 B=2.3

b) 

Fig. 41 Cp results with various VGJ spacing, low TI, Re=50,000: a) Cp profile, b) Integrated 

Cp. 

NASA/CR—2012-217415 288
−1.5 −1 −0.5 0 0.5 1 1.5

0

0.5

1

1.5

φ/Lφ

ψ

B3

B4 B5 B6

No jets

L

p

/d=21.3 z/L

p

=0

L

p

/d=21.3 z/L

p

=0.25

L

p

/d=21.3 z/L

p

=0.5

L

p

/d=21.3 z/L

p

=0.75

L

p

/d=21.3 z/L

p

=1.0

Fig. 42 Ψ profiles, low TI, Re=50,000, steady VJGs, B=2.0, Lp/d=21.3. 

NASA/CR—2012-217415 289
−1.5 −1 −0.5 0 0.5 1 1.5

0

0.5

1

1.5

φ/Lφ

ψ

B3

B4 B5 B6

No jets

L

p

/d=21.3 z/L

p

=0

L

p

/d=21.3 z/L

p

=0.25

L

p

/d=21.3 z/L

p

=0.5

L

p

/d=21.3 z/L

p

=0.75

L

p

/d=21.3 z/L

p

=1.0

Fig. 43 Ψ profiles, low TI, Re=50,000, pulsed VJGs, F=0.56, D=10%, B=1.0, Lp/d=21.3. 

NASA/CR—2012-217415 290
−1.5 −1 −0.5 0 0.5 1 1.5

0

0.5

1

1.5

φ/Lφ

ψ

B3

B4 B5 B6

No jets

L

p

/d=42.7 z/L

p

=0

L

p

/d=42.7 z/L

p

=0.25

L

p

/d=42.7 z/L

p

=0.5

L

p

/d=42.7 z/L

p

=0.75

L

p

/d=42.7 z/L

p

=1.0

Fig. 44 Ψ profiles, low TI, Re=50,000, steady VJGs, B=2.0, Lp/d=42.7. 

NASA/CR—2012-217415 291
−1.5 −1 −0.5 0 0.5 1 1.5

0

0.5

1

1.5

φ/Lφ

ψ

B3

B4 B5 B6

No jets

L

p

/d=42.7 z/L

p

=0

L

p

/d=42.7 z/L

p

=0.25

L

p

/d=42.7 z/L

p

=0.5

L

p

/d=42.7 z/L

p

=0.75

L

p

/d=42.7 z/L

p

=1.0

Fig. 45 Ψ profiles, low TI, Re=50,000, pulsed VJGs, F=0.56, D=10%, B=1.0, Lp/d=42.7. 

NASA/CR—2012-217415 292
−1.5 −1 −0.5 0 0.5 1 1.5

0

0.5

1

1.5

φ/Lφ

ψ

B3

B4 B5 B6

No jets

L

p

/d=42.7 F=0.56 D=10 B=1

L

p

/d=42.7 F=0.56 D=10 B=1.5

L

p

/d=85.3 F=0.56 D=10 B=2.3

Fig. 46 Ψ profiles, low TI, Re=50,000, pulsed VJGs, F=0.56, D=10%, B=2.0, Lp/d=42.7 and 

85.3, z/Lp=0, various B. 

NASA/CR—2012-217415 293
0 0.05 0.1 0.15 0.2 0.25

0

0.2

0.4

0.6

0.81

y/L

s

U/U

e

z/L

p

=0

z/L

p

=0.25

z/L

p

=0.5

z/L

p

=0.75

z/L

p

=1.0

a) 

NASA/CR—2012-217415 294
0 0.05 0.1 0.15 0.2 0.25

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

0.2

y/Ls

u′/Ue

z/L

p

=0

z/L

p

=0.25

z/L

p

=0.5

z/L

p

=0.75

z/L

p

=1.0

b) 

Fig. 47 Time averaged velocity profiles, low TI, Re=50,000, steady VGJs, B=2.0, Lp/d=21.3: a) 

mean, b) rms. 

NASA/CR—2012-217415 295
0 0.05 0.1 0.15 0.2 0.25

0

0.2

0.4

0.6

0.81

y/L

s

U/U

e

z/L

p

=0

z/L

p

=0.25

z/L

p

=0.5

z/L

p

=0.75

z/L

p

=1.0

a) 

NASA/CR—2012-217415 296
0 0.05 0.1 0.15 0.2 0.25

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

0.2

y/Ls

u′/Ue

z/L

p

=0

z/L

p

=0.25

z/L

p

=0.5

z/L

p

=0.75

z/L

p

=1.0

b) 

Fig. 48 Time averaged velocity profiles, low TI, Re=50,000, pulsed VGJs, F=0.56, D=10%, 

B=1.0, Lp/d=21.3: a) mean, b) rms. 

NASA/CR—2012-217415 297
0 0.05 0.1 0.15 0.2 0.25

0

0.2

0.4

0.6

0.81

y/L

s

U/U

e

z/L

p

=0

z/L

p

=0.25

z/L

p

=0.5

z/L

p

=0.75

z/L

p

=1.0

a) 

NASA/CR—2012-217415 298
0 0.05 0.1 0.15 0.2 0.25

0

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

y/Ls

u′/Ue

z/L

p

=0

z/L

p

=0.25

z/L

p

=0.5

z/L

p

=0.75

z/L

p

=1.0

b) 

Fig. 49 Time averaged velocity profiles, low TI, Re=50,000, steady VGJs, B=2.0, Lp/d=42.7: a) 

mean, b) rms. 

NASA/CR—2012-217415 299
0 0.05 0.1 0.15 0.2 0.25

0

0.2

0.4

0.6

0.81

y/L

s

U/U

e

z/L

p

=0

z/L

p

=0.25

z/L

p

=0.5

z/L

p

=0.75

z/L

p

=1.0

z/L

p

=0, B=1.5

a) 

NASA/CR—2012-217415 300
0 0.05 0.1 0.15 0.2 0.25

0

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

y/Ls

u′/Ue

z/L

p

=0

z/L

p

=0.25

z/L

p

=0.5

z/L

p

=0.75

z/L

p

=1.0

z/L

p

=0, B=1.5

b) 

Fig. 50 Time averaged velocity profiles, low TI, Re=50,000, pulsed VGJs, F=0.56, D=10%, 

B=1.0 and 1.5, Lp/d=42.7: a) mean, b) rms. 

NASA/CR—2012-217415 301
0 0.05 0.1 0.15 0.2 0.25

0

0.2

0.4

0.6

0.81

y/L

s

U/U

e

z/L

p

=0

z/L

p

=0.25

z/L

p

=0.5

z/L

p

=0.75

z/L

p

=1.0

a) 

NASA/CR—2012-217415 302
0 0.05 0.1 0.15 0.2 0.25

0

0.05

0.1

0.15

0.2

0.25

0.3

0.35

0.4

y/Ls

u′/Ue

z/L

p

=0

z/L

p

=0.25

z/L

p

=0.5

z/L

p

=0.75

z/L

p

=1.0

b) 

Fig. 51 Time averaged velocity profiles, low TI, Re=50,000, pulsed VGJs, F=0.56, D=10%, 

B=2.3, Lp/d=85.3: a) mean, b) rms. 

NASA/CR—2012-217415 303
0 0.2 0.4 0.6 0.8

1

0

0.51

1.52

2.5

s/L

s

Cp

No jets

Steady,B=1.0

Steady,B=2.0

Steady,B=2.5

Steady,B=3.0

Inviscid

a) 

NASA/CR—2012-217415 304
0 0.5 1 1.5 2 2.5 3

0.74

0.76

0.78

0.8

0.82

0.84

0.86

0.88

0.9

0.92

B

Cp/Cpinv

b) 

Fig. 52 Cp results, high TI, Re=25,000, steady VGJs: a) Cp profile, b) Integrated Cp. 

NASA/CR—2012-217415 305
0 0.2 0.4 0.6 0.8

1

0

0.51

1.52

2.5

s/L

s

Cp

No jets

10%,B=2.0

10%,B=3.0

50%,B=1.0

50%,B=2.0

50%,B=3.0

Inviscid

a) 

NASA/CR—2012-217415 306
0 0.5 1 1.5 2 2.5 3

0.74

0.75

0.76

0.77

0.78

0.79

0.8

0.81

0.82

0.83

0.84

B

Cp/Cpinv

No jets

Pulsed jets, D=10

Pulsed jets, D=50

b) 

Fig. 53 Cp results, high TI, Re=25,000, pulsed VGJs, F=0.14: a) Cp profile, b) Integrated Cp. 

NASA/CR—2012-217415 307
−1.5 −1 −0.5 0 0.5 1 1.5

0

0.5

1

1.5

φ/Lφ

ψ

No control

10%,B=2.0

10%,B=2.5

10%,B=3.0

50%,B=2.0

50%,B=3.0

Steady,B=3.0

a) 

NASA/CR—2012-217415 308
0 0.5 1 1.5 2 2.5 3

0.35

0.4

0.45

0.5

0.55

B

ψint

No jets

Steady jets

Pulsed jets, D=10

Pulsed jets, D=50

b) 

Fig. 54 Ψ results, high TI, Re=25,000, pulsed VGJs, F=0.14: a) Ψ profile, b) Integrated Ψ. 

NASA/CR—2012-217415 309
0 0.2 0.4 0.6 0.8

1

0

0.51

1.52

2.5

s/L

s

Cp

No jets

10%,B=1.0

10%,B=1.5

10%,B=2.0

10%,B=3.0

50%,B=1.0

50%,B=2.0

50%,B=2.5

50%,B=3.0

Inviscid

a) 

NASA/CR—2012-217415 310
0 0.5 1 1.5 2 2.5 3

0.75

0.8

0.85

0.9

0.95

1

B

Cp/Cpinv

No jets

Pulsed jets, D=10

Pulsed jets, D=50

b) 

Fig. 55 Cp results, high TI, Re=25,000, pulsed VGJs, F=0.28: a) Cp profile, b) Integrated Cp. 

NASA/CR—2012-217415 311
−1.5 −1 −0.5 0 0.5 1 1.5

0

0.5

1

1.5

φ/Lφ

ψ

No control

10%,B=1.5

10%,B=2.0

10%,B=3.0

50%,B=2.0

50%,B=3.0

a) 

NASA/CR—2012-217415 312
0 0.5 1 1.5 2 2.5 3

0.2

0.25

0.3

0.35

0.4

0.45

0.5

0.55

B

ψint

No jets

Pulsed jets, D=10

Pulsed jets, D=50

b) 

Fig. 56 Ψ results, high TI, Re=25,000, pulsed VGJs, F=0.28: a) Ψ profile, b) Integrated Ψ. 

NASA/CR—2012-217415 313
0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 57 Time averaged velocity profiles, high TI, Re=25,000, no VGJs: a) mean, b) rms. 

0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 58 Time averaged velocity profiles, high TI, Re=25,000, pulsed VGJs, F=0.28, D=10%, 

B=2.0: a) mean, b) rms. 

NASA/CR—2012-217415 314
0 0.1 0.2

0.042

High TI, Re=25,000, f=06, d=10, b=2.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

a) 

NASA/CR—2012-217415 315
0 0.1 0.2

0.042

High TI, Re=25,000, f=06, d=10, b=2.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/Ue

b) 

Fig. 59 Phase averaged velocity profiles, high TI, Re=25,000, pulsed VGJs, F=0.28, D=10%, 

B=2.0: a) mean, b) rms. 

NASA/CR—2012-217415 316
0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 60 Time averaged velocity profiles, high TI, Re=25,000, pulsed VGJs, F=0.28, D=10%, 

B=3.0: a) mean, b) rms. 

NASA/CR—2012-217415 317
0 0.1 0.2

0.042

High TI, Re=25,000, f=06, d=10, b=3.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

a) 

NASA/CR—2012-217415 318
0 0.1 0.2

0.042

High TI, Re=25,000, f=06, d=10, b=3.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/Ue

b) 

Fig. 61 Phase averaged velocity profiles, high TI, Re=25,000, pulsed VGJs, F=0.28, D=10%, 

B=3.0: a) mean, b) rms. 

NASA/CR—2012-217415 319
0 0.2 0.4 0.6 0.8

1

0

0.51

1.52

2.5

s/L

s

Cp

No jets

10%,B=0.5

10%,B=0.75

10%,B=1.0

10%,B=1.5

10%,B=2.0

10%,B=2.5

10%,B=3.0

50%,B=0.5

50%,B=1.0

50%,B=2.0

Inviscid

a) 

NASA/CR—2012-217415 320
0 0.5 1 1.5 2 2.5 3

0.8

0.85

0.9

0.95

1

1.05

1.1

1.15

B

Cp/Cpinv

No jets

Pulsed jets, D=10

Pulsed jets, D=50

b) 

Fig. 62 Cp results, high TI, Re=25,000, pulsed VGJs, F=0.56: a) Cp profile, b) Integrated Cp. 

NASA/CR—2012-217415 321
0 0.2 0.4 0.6 0.8

1

0

0.51

1.52

2.5

s/L

s

Cp

No jets

10%,B=0.5

10%,B=0.75

10%,B=1.0

10%,B=1.5

10%,B=2.0

Inviscid

a) 

NASA/CR—2012-217415 322
0 0.5 1 1.5 2

0.8

0.85

0.9

0.95

1

1.05

B

Cp/Cpinv

Steady

F=1.12 D=10

b) 

Fig. 63 Cp results, high TI, Re=25,000, pulsed VGJs, F=1.12: a) Cp profile, b) Integrated Cp. 

NASA/CR—2012-217415 323
−1.5 −1 −0.5 0 0.5 1 1.5

0

0.5

1

1.5

φ/Lφ

ψ

No control

10%,B=0.75

10%,B=1.0

10%,B=2.0

50%,B=0.75

50%,B=1.0

50%,B=2.0

10%,B=1.0,f=24hz

10%,B=2.0,f=24hz

a) 

NASA/CR—2012-217415 324
0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 2

0.25

0.3

0.35

0.4

0.45

0.5

0.55

0.6

B

ψint

No jets

F=0.56, 10%

F=0.56, 50%

F=1.12, 10%

b) 

Fig. 64 Ψ results, high TI, Re=25,000, pulsed VGJs, F=0.56 and 1.12: a) Ψ profile, b) Integrated 

Ψ. 

NASA/CR—2012-217415 325
0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 65 Time averaged velocity profiles, high TI, Re=50,000, pulsed VGJs, F=0.56, D=10%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 326
0 0.1 0.2

0.042

High TI, Re=25,000, f=12, d=10, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

a) 

NASA/CR—2012-217415 327
0 0.1 0.2

0.042

High TI, Re=25,000, f=12, d=10, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/Ue

b) 

Fig. 66 Phase averaged velocity profiles, high TI, Re=50,000, pulsed VGJs, F=0.56, D=10%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 328
0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 67 Time averaged velocity profiles, high TI, Re=50,000, pulsed VGJs, F=0.56, D=10%, 

B=2.0: a) mean, b) rms. 

NASA/CR—2012-217415 329
0 0.1 0.2

0.042

High TI, Re=25,000, f=12, d=10, b=2.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

a) 

NASA/CR—2012-217415 330
0 0.1 0.2

0.042

High TI, Re=25,000, f=12, d=10, b=2.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/Ue

b) 

Fig. 68 Phase averaged velocity profiles, high TI, Re=50,000, pulsed VGJs, F=0.56, D=10%, 

B=2.0: a) mean, b) rms. 

NASA/CR—2012-217415 331
0 0.2 0.4 0.6 0.8

1

0

0.51

1.52

2.5

s/L

s

Cp

No jets

Steady,B=0.5

Steady,B=0.75

Steady,B=1.0

Steady,B=1.5

Steady,B=2.0

Inviscid

a) 

NASA/CR—2012-217415 332
0 0.5 1 1.5 2

0.82

0.84

0.86

0.88

0.9

0.92

0.94

0.96

0.98

1

B

Cp/Cpinv

b) 

Fig. 69 Cp results, high TI, Re=50,000, steady VGJs: a) Cp profile, b) Integrated Cp. 

NASA/CR—2012-217415 333
−1.5 −1 −0.5 0 0.5 1 1.5

0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

No control

Steady,B=1.0

Steady,B=1.5

Steady,B=2.0

a) 

NASA/CR—2012-217415 334
0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 2

0.2

0.25

0.3

0.35

0.4

0.45

B

ψint

b) 

Fig. 70 Ψ results, high TI, Re=50,000, steady VGJs: a) Ψ profile, b) Integrated Ψ. 

NASA/CR—2012-217415 335
0 0.2 0.4 0.6 0.8

1

0

0.51

1.52

2.5

s/L

s

Cp

No jets

10%,B=0.25

10%,B=0.5

10%,B=0.75

10%,B=1.0

10%,B=1.5

10%,B=2.0

50%,B=0.25

50%,B=0.5

50%,B=0.75

50%,B=1.0

50%,B=1.5

50%,B=2.0

Inviscid

a) 

NASA/CR—2012-217415 336
0 0.5 1 1.5 2

0.82

0.84

0.86

0.88

0.9

0.92

0.94

0.96

0.98

1

B

Cp/Cpinv

No jets

Pulsed jets, D=10

Pulsed jets, D=50

b) 

Fig. 71 Cp results, high TI, Re=50,000, pulsed VGJs, F=0.14: a) Cp profile, b) Integrated Cp. 

NASA/CR—2012-217415 337
−1.5

−1−0.50 0.51 1.5

0

0.2

0.4

0.6

0.81

1.2

φ/L

φ

ψ

No control

10%,B=1.0

10%,B=1.5

10%,B=2.0

50%,B=1.0

50%,B=1.5

a) 

NASA/CR—2012-217415 338
0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 2

0.2

0.25

0.3

0.35

0.4

0.45

0.5

B

ψint

No jets

Pulsed jets, D=10

Pulsed jets, D=50

b) 

Fig. 72 Ψ results, high TI, Re=50,000, pulsed VGJs, F=0.14: a) Ψ profile, b) Integrated Ψ. 

NASA/CR—2012-217415 339
0 0.2 0.4 0.6 0.8

1

0

0.51

1.52

2.5

s/L

s

Cp

No jets

10%,B=0.25

10%,B=0.5

10%,B=0.75

10%,B=1.0

10%,B=1.5

50%,B=0.25

50%,B=0.5

50%,B=0.75

50%,B=1.0

Inviscid

a) 

NASA/CR—2012-217415 340
0 0.5 1 1.5

0.82

0.84

0.86

0.88

0.9

0.92

0.94

0.96

0.98

1

B

Cp/Cpinv

No jets

Pulsed jets, D=10

Pulsed jets, D=50

b) 

Fig. 73 Cp results, high TI, Re=50,000, pulsed VGJs, F=0.28: a) Cp profile, b) Integrated Cp. 

NASA/CR—2012-217415 341
−1.5

−1−0.50 0.51 1.5

0

0.2

0.4

0.6

0.81

1.2

φ/L

φ

ψ

No control

10%,B=0.75

10%,B=1.0

50%,B=0.5

50%,B=0.75

50%,B=1.0

a) 

NASA/CR—2012-217415 342
0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

0.2

0.25

0.3

0.35

0.4

0.45

0.5

B

ψint

No jets

Pulsed jets, D=10

Pulsed jets, D=50

b) 

Fig. 74 Ψ results, high TI, Re=50,000, pulsed VGJs, F=0.28: a) Ψ profile, b) Integrated Ψ. 

NASA/CR—2012-217415 343
0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 75 Time averaged velocity profiles, high TI, Re=50,000, no VGJs: a) mean, b) rms. 

0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 76 Time averaged velocity profiles, high TI, Re=50,000, pulsed VGJs, F=0.28, D=10%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 344
0 0.1 0.2

0.042

High TI, Re=50,000, f=12, d=10, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

a) 

NASA/CR—2012-217415 345
0 0.1 0.2

0.042

High TI, Re=50,000, f=12, d=10, b=1.0

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/Ue

b) 

Fig. 77 Phase averaged velocity profiles, high TI, Re=50,000, pulsed VGJs, F=0.28, D=10%, 

B=1.0: a) mean, b) rms. 

NASA/CR—2012-217415 346
0 0.2 0.4 0.6 0.8

1

0

0.51

1.52

2.5

s/L

s

Cp

No jets

10%,B=0.25

10%,B=0.5

10%,B=0.75

10%,B=1.0

50%,B=0.25

50%,B=0.5

50%,B=0.75

Inviscid

a) 

NASA/CR—2012-217415 347
0 0.2 0.4 0.6 0.8 1

0.82

0.84

0.86

0.88

0.9

0.92

0.94

0.96

0.98

1

B

Cp/Cpinv

No jets

Pulsed jets, D=10

Pulsed jets, D=50

b) 

Fig. 78 Cp results, high TI, Re=50,000, pulsed VGJs, F=0.56: a) Cp profile, b) Integrated Cp. 

NASA/CR—2012-217415 348
−1.5

−1−0.50 0.51 1.5

0

0.2

0.4

0.6

0.81

1.2

φ/L

φ

ψ

No control

10%,B=0.5

10%,B=0.75

10%,B=1.0

50%,B=0.5

50%,B=0.75

a) 

NASA/CR—2012-217415 349
0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

0.2

0.25

0.3

0.35

0.4

0.45

0.5

B

ψint

No jets

Pulsed jets, D=10

Pulsed jets, D=50

b) 

Fig. 79 Ψ results, high TI, Re=50,000, pulsed VGJs, F=0.56: a) Ψ profile, b) Integrated Ψ. 

NASA/CR—2012-217415 350
0 1 0 1 0 1 0 1 0 1 0 1

0

0.1

0.2

U/U

e

y/Ls

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0

0.1

0.2

u′/U

e

y/Ls

Fig. 80 Time averaged velocity profiles, high TI, Re=50,000, pulsed VGJs, F=0.56, D=10%, 

B=0.75: a) mean, b) rms. 

NASA/CR—2012-217415 351
0 0.1 0.2

0.042

High TI, Re=50,000, f=24, d=10, b=0.75

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

a) 

NASA/CR—2012-217415 352
0 0.1 0.2

0.042

High TI, Re=50,000, f=24, d=10, b=0.75

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/Ue

b) 

Fig. 81 Phase averaged velocity profiles, high TI, Re=50,000, pulsed VGJs, F=0.56, D=10%, 

B=0.75: a) mean, b) rms. 

NASA/CR—2012-217415 353
B

F

0 0.51 1.52 2.53

0

0.2

0.4

0.6

0.81

0.2

0.25

0.3

0.35

0.4

0.45

0.5

0.55

a) 

NASA/CR—2012-217415 354
B

F

0 0.51 1.52 2.53

0

0.2

0.4

0.6

0.81

0.2

0.25

0.3

0.35

0.4

0.45

0.5

0.55

b) 

NASA/CR—2012-217415 355
B

F

0 0.51 1.52 2.53

0

0.2

0.4

0.6

0.81

0.2

0.25

0.3

0.35

0.4

0.45

0.5

0.55

c) 

NASA/CR—2012-217415 356
B

F

0 0.5 1 1.5 2 2.5 3

0

0.2

0.4

0.6

0.8

1

0.2

0.25

0.3

0.35

0.4

0.45

0.5

0.55

d) 

Fig. 82 Integrated total pressure loss contours as function of F and B: a) low TI, Re=25,000, b) 

low TI, Re=50,000, c) high TI, Re=25,000, d) high TI, Re=50,000. 

NASA/CR—2012-217415 357
−0.5 −0.4 −0.3 −0.2 −0.1 0 0.1 0.2 0.3 0.4 0.5

0.5

0.6

0.7

0.8

0.9

1

1.1

1.2

1.3

φ/Lφ

U′/U

mean

Rod wake

Cascade wake, Re=50,000, 0.3Cx

Cascade wake, Re=50,000, 0.63Cx

Cascade wake, Re=200,000, 0.3Cx

Cascade wake, Re=200,000, 0.63Cx

a) 

NASA/CR—2012-217415 358
−0.5 −0.4 −0.3 −0.2 −0.1 0 0.1 0.2 0.3 0.4 0.5

0

0.05

0.1

0.15

0.2

0.25

φ/Lφ

u′/U

mean

Rod wake

Cascade wake, Re=50,000, 0.3Cx

Cascade wake, Re=50,000, 0.63Cx

Cascade wake, Re=200,000, 0.3Cx

Cascade wake, Re=200,000, 0.63Cx

b) 

Fig. 83 Velocity in wakes of cascade blades for low TI, no VGJ, no wake cases and of wake 

generator rods: a) mean, b) rms. 

NASA/CR—2012-217415 359
−0.5 −0.4 −0.3 −0.2 −0.1 0 0.1 0.2 0.3 0.4 0.5

0.7

0.75

0.8

0.85

0.9

0.95

1

1.05

1.1

1.15

φ/Lφ

U/U

mean

Rod wake

Cascade wake

a) 

NASA/CR—2012-217415 360
−0.5 −0.4 −0.3 −0.2 −0.1 0 0.1 0.2 0.3 0.4 0.5

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

0.2

φ/Lφ

u′/U

mean

Rod wake

Cascade wake

b) 

Fig. 84 Phase averaged velocity in wakes of cascade blades for low TI, Re=50,000, F=0.56, 

D=10%, B=1.0 case and of wake generator rods: a) mean, b) rms. 

NASA/CR—2012-217415 361
a) 

NASA/CR—2012-217415 362
b) 

Fig. 85 Fourier spectra of wake from cascade blades, no VGJs: a) Re=50,000, b) Re=200,000. 

NASA/CR—2012-217415 363
Fig. 86 Fourier spectra of wake from cascade blades, Re=50,000, pulsed VGJs, F=0.56, D=10%, 

B=1.0. 

NASA/CR—2012-217415 364
−1.5

−1

−0.5

0

0.5

1

1.5

10

−2

10

−1

10

0

10

1

10

2

01234567

x 10

−3

φ/L

φ

F

f⋅PSD/U2

e

a) 

NASA/CR—2012-217415 365
−1.5

−1

−0.5

0

0.5

1

1.5

10

−2

10

−1

10

0

10

1

10

2

0

0.51

1.52

2.53

3.5

x 10

−3

φ/L

φ

F

f⋅PSD/U2

e

b) 

NASA/CR—2012-217415 366
−1.5

−1

−0.5

0

0.5

1

1.5

10

−2

10

−1

10

0

10

1

10

2

0123456

x 10

−3

φ/L

φ

F

f⋅PSD/U2

e

c) 

NASA/CR—2012-217415 367
−1.5

−1

−0.5

0

0.5

1

1.5

10−2

10−1

100

101

102

0

1

2

3

4

5

x 10−3

φ/Lφ

F

f⋅PSD/U2

e

d) 

Fig. 87 Wavelet spectra of wake from cascade blades, Re=50,000, pulsed VGJs, F=0.56, 

D=10%, B=1.0: a) Mexican Hat, t/T=0.833, b) Mexican hat, t/T=0.5, c) Morlet, t/T=0.833, d) 

Morlet, t/T=0.5. 

NASA/CR—2012-217415 368
0

0.2

0.4

0.6

0.8

1

1.2

1.4

10

−2

10

−1

10

0

10

1

10

2

0

0.51

1.52

2.53

3.5

x 10

−3

φ/L

φ

F

f⋅PSD/U2

e

a) 

NASA/CR—2012-217415 369
0

0.2

0.4

0.6

0.8

1

1.2

1.4

10−2

10−1

100

101

102

0

1

2

3

4

5

x 10−3

φ/Lφ

F

f⋅PSD/U2

e

b) 

Fig. 88 Wavelet spectra of wake generator rods: a) Morlet, b) Mexican Hat. 

NASA/CR—2012-217415 370
0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

0

0.5

1

1.5

2

2.5

s/L

s

Cp

Re=25,000, 6hz

Re=37,500, 9hz

Re=50,000, 12hz

Re=100,000, 12hz

Re=200,000, 12hz

Re=25,000, 12hz

Re=50,000, 6hz

Re=200,000, 6 hz

Inviscid

Fig. 89 Cp profiles, low TI, unsteady wakes, no VGJs. 

NASA/CR—2012-217415 371
−1.5 −1 −0.5 0 0.5 1 1.5

0

0.2

0.4

0.6

0.8

1

1.2

φ/Lφ

ψ

B3

B4

B5 B6

Re=25,000, 6Hz

Re=37,500, 9Hz

Re=50,000, 12Hz

Re=100,000, 12Hz

Re=200,000, 12Hz

Re=25,000, 12Hz

Re=50,000, 6Hz

Re=200,000, 6Hz

Fig. 90 Ψ profiles, low TI, unsteady wakes, no VGJs. 

NASA/CR—2012-217415 372
0 0.1 0.2

0.042

Re=25,000, Wake 6 Hz

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

a) 

NASA/CR—2012-217415 373
0 0.1 0.2

0.042

Re=25,000, Wake 6 Hz

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/Ue

b) 

Fig. 91 Phase averaged velocity profiles, low TI, Re=25,000, F=0.28 wakes: a) mean, b) rms. 

NASA/CR—2012-217415 374
0 0.1 0.2

0.042

Re=37,500, Wake 9 Hz

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

a) 

NASA/CR—2012-217415 375
0 0.1 0.2

0.042

Re=37,500, Wake 9 Hz

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/Ue

b) 

Fig. 92 Phase averaged velocity profiles, low TI, Re=37,500, F=0.28 wakes: a) mean, b) rms. 

NASA/CR—2012-217415 376
0 0.1 0.2

0.042

Re=50,000, Wake 12 Hz

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

a) 

NASA/CR—2012-217415 377
0 0.1 0.2

0.042

Re=50,000, Wake 12 Hz

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/Ue

b) 

Fig. 93 Phase averaged velocity profiles, low TI, Re=50,000, F=0.28 wakes: a) mean, b) rms. 

NASA/CR—2012-217415 378
0 0.1 0.2

0.042

Re=50,000, Wake 6 Hz

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

a) 

NASA/CR—2012-217415 379
0 0.1 0.2

0.042

Re=50,000, Wake 6 Hz

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/Ue

b) 

Fig. 94 Phase averaged velocity profiles, low TI, Re=50,000, F=0.14 wakes: a) mean, b) rms. 

NASA/CR—2012-217415 380
The following figures show phase averaged velocity and turbulence profiles for cases with VGJs and 

wakes with various timings. The names for the timings correspond to the original data file names and not the 

new names used in conference papers. 

NASA/CR—2012-217415 381
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

7/21/2010 Re=25,000 Blue: Low TI Red: High TI Timing a 

NASA/CR—2012-217415 382
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

7/21/2010 Re=25,000 Blue: Low TI Red: High TI Timing b 

NASA/CR—2012-217415 383
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

7/21/2010 Re=25,000 Blue: Low TI Red: High TI Timing c 

NASA/CR—2012-217415 384
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

7/21/2010 Re=25,000 Blue: Low TI Red: High TI Timing d 

NASA/CR—2012-217415 385
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

7/21/2010 Re=25,000 Blue: Low TI Red: High TI Timing e 

NASA/CR—2012-217415 386
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

7/21/2010 Re=25,000 Blue: Low TI Red: High TI Timing f 

NASA/CR—2012-217415 387
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

Re=25,000 Low TI Timing i 

NASA/CR—2012-217415 388
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

Re=25,000 Low TI Timing j 

NASA/CR—2012-217415 389
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

7/21/2010 Re=50,000 Blue: Low TI Red: High TI Timing a 

NASA/CR—2012-217415 390
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

7/21/2010 Re=50,000 Blue: Low TI Red: High TI Timing b 

NASA/CR—2012-217415 391
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

7/21/2010 Re=50,000 Blue: Low TI Red: High TI Timing c 

NASA/CR—2012-217415 392
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

7/21/2010 Re=50,000 Blue: Low TI Red: High TI Timing d 

NASA/CR—2012-217415 393
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

7/21/2010 Re=50,000 Blue: Low TI Red: High TI Timing e 

NASA/CR—2012-217415 394
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

010101010101

0 0.1 0.2

1

U/U

e

Re=50,000 Low TI Timing g 

NASA/CR—2012-217415 395
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/U

e

7/21/2010 Re=25,000 Blue: Low TI Red: High TI Timing a 

NASA/CR—2012-217415 396
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/U

e

7/21/2010 Re=25,000 Blue: Low TI Red: High TI Timing b 

NASA/CR—2012-217415 397
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/U

e

7/21/2010 Re=25,000 Blue: Low TI Red: High TI Timing c 

NASA/CR—2012-217415 398
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/U

e

7/21/2010 Re=25,000 Blue: Low TI Red: High TI Timing d 

NASA/CR—2012-217415 399
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/U

e

7/21/2010 Re=25,000 Blue: Low TI Red: High TI Timing e 

NASA/CR—2012-217415 400
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/U

e

7/21/2010 Re=25,000 Blue: Low TI Red: High TI Timing f 

NASA/CR—2012-217415 401
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.20 0.20 0.20 0.20 0.20 0.2

0 0.1 0.2

1

u′/U

e

Re=25,000 Low TI Timing i 

NASA/CR—2012-217415 402
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.20 0.20 0.20 0.20 0.20 0.2

0 0.1 0.2

1

u′/U

e

Re=25,000 Low TI Timing j 

NASA/CR—2012-217415 403
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/U

e

7/21/2010 Re=50,000 Blue: Low TI Red: High TI Timing a 

NASA/CR—2012-217415 404
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/U

e

7/21/2010 Re=50,000 Blue: Low TI Red: High TI Timing b 

NASA/CR—2012-217415 405
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/U

e

7/21/2010 Re=50,000 Blue: Low TI Red: High TI Timing c 

NASA/CR—2012-217415 406
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/U

e

7/21/2010 Re=50,000 Blue: Low TI Red: High TI Timing d 

NASA/CR—2012-217415 407
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.2 0 0.2 0 0.2 0 0.2 0 0.2 0 0.2

0 0.1 0.2

1

u′/U

e

7/21/2010 Re=50,000 Blue: Low TI Red: High TI Timing e 

NASA/CR—2012-217415 408
0 0.1 0.2

0.042

0 0.1 0.2

0.083

0 0.1 0.2

0.125

0 0.1 0.2

0.167

0 0.1 0.2

0.208

0 0.1 0.2

0.25

0 0.1 0.2

0.292

0 0.1 0.2

0.333

0 0.1 0.2

0.375

0 0.1 0.2

0.417

0 0.1 0.2

0.458

0 0.1 0.2

y/L

0.5

s

0 0.1 0.2

0.542

0 0.1 0.2

0.583

0 0.1 0.2

0.625

0 0.1 0.2

0.667

0 0.1 0.2

0.708

0 0.1 0.2

0.75

0 0.1 0.2

0.792

0 0.1 0.2

0.833

0 0.1 0.2

0.875

0 0.1 0.2

0.917

0 0.1 0.2

0.958

0 0.20 0.20 0.20 0.20 0.20 0.2

0 0.1 0.2

1

u′/U

e

Re=50,000 Low TI Timing g 

NASA/CR—2012-217415 409
The following figures show time space plots of phase averaged velocity and turbulence in the wake 0.6 Cx

downstream of the trailing edges for cases with VGJs and wakes with various timings. The names for the 

timings correspond to the original data file names and not the new names used in conference papers. 

NASA/CR—2012-217415 410
φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=25,000, Low TI, Timing a, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=25,000, Low TI, Timing b, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=25,000, Low TI, Timing c, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=25,000, Low TI, Timing a, u′/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=25,000, Low TI, Timing b, u′/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=25,000, Low TI, Timing c, u′/Ue

NASA/CR—2012-217415 411
φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=25,000, Low TI, Timing d, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=25,000, Low TI, Timing e, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=25,000, Low TI, Timing f, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=25,000, Low TI, Timing d, u′/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=25,000, Low TI, Timing e, u′/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=25,000, Low TI, Timing f, u′/Ue

NASA/CR—2012-217415 412
φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=25,000, Low TI, Timing i, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=25,000, Low TI, Timing j, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=25,000, Low TI, Timing i, u′/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=25,000, Low TI, Timing j, u′/Ue

NASA/CR—2012-217415 413
φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=50,000, Low TI, Timing a, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=50,000, Low TI, Timing b, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=50,000, Low TI, Timing c, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=50,000, Low TI, Timing a, u′/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=50,000, Low TI, Timing b, u′/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=50,000, Low TI, Timing c, u′/Ue

NASA/CR—2012-217415 414
φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=50,000, Low TI, Timing d, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=50,000, Low TI, Timing e, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=50,000, Low TI, Timing g, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=50,000, Low TI, Timing d, u′/Ue

−1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=50,000, Low TI, Timing e, u′/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=50,000, Low TI, Timing g, u′/Ue

NASA/CR—2012-217415 415
φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=25,000, High TI, Timing a, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=25,000, High TI, Timing b, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=25,000, High TI, Timing c, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=25,000, High TI, Timing a, u′/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=25,000, High TI, Timing b, u′/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=25,000, High TI, Timing c, u′/Ue

NASA/CR—2012-217415 416
φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=25,000, High TI, Timing d, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=25,000, High TI, Timing e, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=25,000, High TI, Timing f, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=25,000, High TI, Timing d, u′/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=25,000, High TI, Timing e, u′/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=25,000, High TI, Timing f, u′/Ue

NASA/CR—2012-217415 417
φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=50,000, High TI, Timing a, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=50,000, High TI, Timing b, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=50,000, High TI, Timing c, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=50,000, High TI, Timing a, u′/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=50,000, High TI, Timing b, u′/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=50,000, High TI, Timing c, u′/Ue

NASA/CR—2012-217415 418
φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=50,000, High TI, Timing d, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.7

0.8

0.9

1

1.1

1.2

Re=50,000, High TI, Timing e, Mean U/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=50,000, High TI, Timing d, u′/Ue

φ/Lφ

t/T −1.5 −1 −0.5 0 0.5 1 1.5

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

0.18

Re=50,000, High TI, Timing e, u′/Ue

NASA/CR—2012-217415 419

Appendix C 

Experimental Data Archive 

A supplementary CD–ROM is available containing data files with processed experimental data from 

the study and text files with explanations of the file formats for the data. 

The complete PDF file of Part I of this report (NASA/CR-2012-217415) along with Part II 

(NASA/CR-2012-217416) and Part III (NASA/CR-2012-217417), are also on the CD–ROM. The 

CD–ROM can be obtained from the Center for AeroSpace Information (CASI) Web site at 

http://www.sti.nasa.gov. 

NASA/CR—2012-217415 421
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4. TITLE AND SUBTITLE

Flow Control Under Low-Pressure Turbine Conditions Using Pulsed Jets 

Final Report

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NNC07IA10I; N00189-07-P-A253 

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Volino, Ralph, J.; Ibrahim, Mounir, B. 

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WBS 561581.02.08.03.47.02.01 

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13. SUPPLEMENTARY NOTES

This report contains a supplemental CD-ROM containing detailed data along with the full text PDF files of this report and NASA/CR-2012-

217416 and NASA/CR-2012-217417. The CD-ROM can be obtained from the Center for AeroSpace Information (CASI) Web site at 

http://www.sti.nasa.gov 

Grant technical monitors, Anthony J. Strazisar, Office of the Chief Scientist, Glenn Research Center, organization code ASOO, James D. 

Heidmann, Aeropropulsion Division, Glenn Research Center, organization code RTTO, David E. Ashpis, Aeropropulsion Division, Glenn 

Research Center, organization code RTTO, ashpis@nasa.gov 

14. ABSTRACT

This publication is the final report of research performed under an NRA/Cooperative Interagency Agreement, and includes a supplemental CD-ROM with detailed data. It is 

complemented by NASA/CR-2012-217416 and NASA/CR-2012-217417 which include a Ph.D. Dissertation and an M.S. thesis respectively, performed under this contract. In 

this study the effects of unsteady wakes and flow control using vortex generator jets (VGJs) were studied experimentally and computationally on the flow over the L1A low 

pressure turbine (LPT) airfoil. The experimental facility was a six passage linear cascade in a low speed wind tunnel at the U.S. Naval Academy. In parallel, computational 

work using the commercial code FLUENT (ANSYS, Inc.) was performed at Cleveland State University, using Unsteady Reynolds Averaged Navier Stokes (URANS) and 

Large Eddy Simulations (LES) methods. In the first phase of the work, the baseline flow was documented under steady inflow conditions without flow control. URANS 

calculations were done using a variety of turbulence models. In the second phase of the work, flow control was added using steady and pulsed vortex generator jets. The VGJs 

successfully suppressed separation and reduced aerodynamic losses. Pulsed operation was more effective and mass flow requirements are very low. Numerical simulations of 

the VGJs cases showed that URANS failed to capture the effect of the jets. LES results were generally better. In the third phase, effects of unsteady wakes were studied. 

Computations with URANS and LES captured the wake effect and generally predicted separation and reattachment to match the experiments. Quantitatively the results were 

mixed. In the final phase of the study, wakes and VGJs were combined and synchronized using various timing schemes. The timing of the jets with respect to the wakes had 

some effect, but in general once the disturbance frequency was high enough to control separation, the timing was not very important.

15. SUBJECT TERMS

Gas turbines; Tubomachinery; Turbine; Low pressure turbine; Turbulence; Flow control; Synthetic jets; CFD; LES; Turbulence 

models; Wakes; Separation; Vortex generated jets 

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