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EXPERIMENTAL AND COMPUTATIONAL STUDY OF FLOW IMPROVEMENT THROUGH SIGMOID AIR INTAKE DUCTS

USING FLOW DEFLECTOR

Akshoy Ranjan Paul1 and Kalyan Kuppa.2

Motilal Nehru National Institute of Technology, Allahabad, Uttar Pradesh, 211 004, India

Navanshu Tripathi3

Motilal Nehru National Institute of Technology, Allahabad, Uttar Pradesh, 211 004, India

and

Tanmay Rajpathak4

University of Florida, Gainesville, Florida, 32611, USA

Rapid expansion and advancement in military aircraft industry demands light weight, compact and efficient gas turbine engines, which may partly be achieved by the induction of sigmoid shaped air-intake ducts (S-shaped diffusing ducts). These are used to decelerate the flow within a limited space, maintaining a uniform total pressure distribution at its exit. But due to curvature with inflexion, the flow development within such duct is complicated and flow parameters are not uniform at the exit. In the present study, a twin bladed flow deflector has been designed and installed at the duct's inlet in order to uniform the flow pattern at its exit. The paper presents a detailed experimental and numerical investigation of flow through a 30 /30

diffusing duct using a flow deflector. A five-hole pressure probe has been employed to carry out detailed flow study at the exit plane of the duct to know the total flow characteristics. Considering the computational time required and the type of numerical flow solver (FLUENT 6.2) available for the study, the k e realizable turbulence model with enhanced wall treatment has been used. Flow pattern at the exit is appeared to be more uniform with the installation of flow deflector as revealed from the figures of mean velocity and total pressure contours.

Nomenclature

0 , sA A = model constants, dimensionless

rA = area ratio, dimensionless

1 2,C C

= model constants, dimensionless

1 3, ,C C Cε ε µ

= model constants, dimensionless

iD = inlet diameter, mm

bG

= generation of turbulence kinetic energy due to buoyancy, N m−2 s−1

kG

= generation of turbulence kinetic energy due to mean velocity gradients, N m−2 s−1

k = turbulent kinetic energy, m2 s−2

L

= centerline length, mm

1 Lecturer, Department of Applied Mechanics, MNNIT,Allahabad, India . e-mail: arpaul2k@yahoo.co.in 2 Graduate Student, Department of Mechanical Engineering, MNNIT, Allahabad, India. 3 Graduate Student, Department of Mechanical Engineering, MNNIT, Allahabad, India. 4 Graduate Student, Department of Mechanical Engineering, University of Florida, Gainesville, USA.

26th AIAA Applied Aerodynamics Conference18 - 21 August 2008, Honolulu, Hawaii

AIAA 2008-7513

Copyright © 2008 by the American Institute of Aeronautics and Astronautics, Inc. All rights reserved.

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m = air mass, kg

dp

= dynamic pressure at inlet, N m−2

cR = radius of curvature, mm Re

= Reynolds number at inlet, dimensionless

S = modulus of the mean rate of strain tensor

ijS = mean strain rate

, kS Sε = user-defined source terms U

= mass averaged inlet velocity, m s−1

u = velocity, m s−1

, ,x y zu u u

= velocity components, m s−1

u

= time averaged velocity, m s−1

x = longitudinal direction

ix = tensorial notation of Cartesian coordinate system

MY

= contribution of fluctuation dilatation in the compressible turbulence to the overall dissipation rate, N m−2 s−1

Subscripts ,i j

= tensorial notations Greek letters

β∆

= centerline curvature, degree ε

= turbulent kinetic energy dissipation rate, m2 s−3

φ

= scalar quantity, such as pressure, energy or species concentration η

= nondimensional parameter, Sk ε

µ

= dynamic viscosity coefficient, N s m−2

ν

= molecular kinematic viscosity coefficient, m2 s−1

tν

= turbulent or eddy kinematic viscosity coefficient ρ

= density of fluid, kg m−3

, kεσ σ

= turbulent Prandtl number for ε and k, dimensionless

ijΩ

= mean rate-of-rotation tensor viewed in a rotating reference frame with the angular velocity kω

kω

=angular velocity, rad s−1

I. Introduction HE main purpose of the S-shaped (sigmoid) diffusing duct is to deliver airflow from the wing or fuselage intake to the engine compressor. A compressible, subsonic inlet flow conditions usually exist for this

application. A secondary purpose of an S-duct is to cover up the compressor face and provide a multiple bounce cavity for radar reflection. A few examples of commercial aircraft with S-ducts include the Boeing-727 and Lockheed Tristar L-1011. Besides, military aircraft like General Dynamics F-16, McDonnell Douglas F-18, and the Lockheeed YF-22 also use S-shaped ducts.

These intake ducts have to meet the varying flow demands of the engines because these aircrafts have to perform highly complicated maneuvers. The current experience shows that flow pattern (mainly total pressure distribution) is not uniform throughout the sigmoid duct, especially at the exit plane of the duct. Moreover, the exit flow of the diffusers gets distorted due to strong secondary flow and swirl present inside the duct, which is complicated due to presence of flow in boundary layer separation inside the diffuser. This is attributed to the combined effect of twin bends, inflexion and diffusion in the downstream. Therefore, an uneven impact load is experienced at the downstream components of the sigmoid duct like air compressor, combustor etc., which may lead to catastrophic failure and hence undesirable from the design point of view.

The rapid advancement of modern aircraft engines necessitates the study of sigmoid ducts to improve the velocity distribution and tackle the self-generated swirl at its exit. Guo & Seddon [1] used two methods in order to

T

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reduce the magnitude of swirl at high angle of attack. Among them, one was to change the distribution of pressure by means of a 'spoiler'. Of the 'anti-swirl devices', the 'perforated spoiler' is the more powerful and can be sized either to reverse the swirl direction or to eliminate the swirl completely.

A number of vortex control device [2], vortex reduced device [3] and automatic adjustable blade [4] were installed near the inlet of the sigmoid duct of rectangular as well as circular cross-section to curb the swirl at the exit of the diffuser. All these methods were proven effective for measuring swirl and improving flow field in sigmoid ducts. Reichert & Wendt [5, 6] used vortex generators and eight-bladed stationary ‘pinwheel device’ at the upstream

of round sigmoid duct ( rA = 1.5, β∆ = 30°, cR = 1021 mm.) in order to reduce total pressure non-uniformity at its exit. Flow visualization achieved by fluorescent oil-dot technique revealed the improved flow pattern at the exit of the duct. From the experiments, they concluded that most of the degraded aerodynamic performance results from secondary flow indicated by streamline curvature producing a pair of naturally occurring vortices. In the second case, the ingested vortex reduced the flow separation inside the duct and promoted stronger cross-flow, which leads to more uniform pressure distribution at the exit of the duct. In another study, Anderson & Gibb [7] showed reductions in steady state as well as dynamic distortion at the engine face of up to 80% in the case of circular

sigmoid duct ( rA = 2.0) by installing a vortex generator. They clarified the fact that the vortex generator re-structured the secondary low to minimize engine face distortion.

Dominy et.al. [8] carried out experiment in an S-shaped annular duct with 34 inlet swirl vanes and reported that the generation of swirl and wakes had a significant effect upon the detailed structure of the flow. Computer simulation in this respect also had a good agreement with the experimental results. Anderson [9] developed an economical robust design methodology for micro-scale secondary flow control in compact inlet diffusers. With this design methodology, the secondary flow control in the exit is possible to a greater extent if it is optimally designed.

Recently, Pradeep [10] and Sullerey [11] carried out performance enhancement of a circular S-diffuser by secondary flow control using vortex generator jets (VGJ). A 20% decrease in total pressure loss and flow distortion coefficients in the diffuser was reported by effectively controlling the secondary flow using VGJ. For serpentine ducts, 35% improvement in average static pressure recovery coefficient using vortex generator jets were reported in addition to a considerable decrease in the outflow swirl and about 10% decrease in the intensity of distortion.

It has been observed from peer review of related literatures that a lot of vortex generators, fences, grills, spoilers, vortex ingestion devices etc. have been installed so far at the upstream of the diffuser to improve the flow field.

Inspired from these studies, an attempt is made to 'uniform' the flow at 30°/30° sigmoid duct ( rA = 2.25, iL D =

6.0, cR = 287.5 mm, Re = 1.5×105) with installation of a 'twin-bladed' flow deflector (refer Fig. 1). The deflector was designed as per Mathur [12]. The blade angle and swirl angle of the deflector are 30° and 45° respectively as seen in Fig.1. In this study, five-hole pressure probe was used to carry out detailed flow study at the exit plane of diffuser to reveal the total flow characteristics.

II. Experimental Set-up and Procedure The experimental set-up consists of a centrifugal blower, a contraction cone, a reducer and the test piece

(sigmoid duct), the details of which is described by Paul et.al. [13]. The sigmoid duct (Fig. 2) designed is based on linear area-ratio from inlet to exit as per the method outlined by Fox & Kline [14] and are made of fiberglass. The inlet diameter of the diffuser is kept as 50 mm. The test diffuser is added with a straight pipe of length twice the inlet

diameter at its exit to minimize the atmospheric effect. The inlet section is chosen at iD upstream of the duct to avoid downstream effects. The tests have been conducted with air as a working fluid. During the experiments, the flow rate has been maintained constant by adjusting the opening of the blower window and simultaneously checking the pressure drop between the inlet and the exit of the contraction cone. During experiments, the Re has been fixed

at Re = 1.5×105 based on iD

which corresponds to the free stream air velocity of 40 m s−1. All the measured

pressures have been normalized with dp

= 896 N m−2.

III. Instrumentation and Measurement Technique Considering the size and geometry of the test diffuser and optimum accuracy of the results, a single point silk

tuft probe has been used for flow visualization study. The study includes the detailed flow visualization with and

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without installing deflector at diffuser inlet. The orientation of deflector has also been varied and hence, the optimum orientation is selected. A microprocessor based digital micromanometer has been used to measure wall static pressures at various locations on both sides of the curved planes (wall-A and wall-B).

In the present experimentation, a hemispherical tip ‘cobra-shaped’ five-hole pressure probe has been used for the measurement of mean velocity, cross-velocity, longitudinal velocity and static & total pressure variation at the exit cross-section of the diffuser. The probe is designed and fabricated according to the criteria outlined by Bryer & Pankhurst [15]. The probe tip is given a ‘cobra’ shape to avoid stem effects at its tip. The ‘cobra’ shape of the sensing head has also been allowed measurements close to the walls of the diffuser. The overall length of the probe is chosen as 250 mm. The free ends of all five hypodermic tubes have been connected to the flexible tubes for the connection with the pressure scanner. The probe is calibrated against a standard pitot tube in a subsonic low-

turbulence wind tunnel at a constant air flow of 40 m s−1 corresponding to Re = 10256, calculated on the basis of the probe tip diameter of 4 mm. The calibration is carried out in yaw-pitch mode, in which the yaw angles of the probe are varied for a particular value of pitch angle. Once, the curves drawn from the calibration data and their respective polynomial expressions are obtained, the probe is then ready for the successful use in unknown flow field within the diffuser passage. A microprocessor controlled traversing mechanism has been used for traversing the probes precisely into the flow field.

IV. Mathematical Formulation and Numerical Technique The experiment performed here has been validated with help of computational fluid dynamics (CFD) technique,

which are described below.

4.1 Governing equations The present simulations have been performed using the general purpose CFD code- ‘FLUENT version 6.2’

[16] based on finite volume approach. The basic governing equations used are as follows (Biswas [17]): Continuity equation:

( )0

i

i

u

x t

ρ ρ∂ ∂+ =∂ ∂

Momentum equation:

( )2 / 3j i j

tj i j i

p kDu u u

Dt x x x xν

∂ + ∂ ∂ ∂= − + + ∂ ∂ ∂ ∂

tν can be expressed empirically as

( )2t C kµν ε=

For closure solution, some additional equations are required to solve for the unknown parameter tν .

4.2 Choice of turbulence model Selection of an appropriate turbulence model is critical for obtaining correct solution. A study conducted by

Kumar et.al. [18] of FLUENT Inc. on a similar S-duct problem has used the k−e

realizable model and Reynolds Stress Model (RSM). The study also reports good agreement with the NASA experimental results. Considering the computational time required and the type of numerical flow solver available for the study, the k−e realizable turbulence model with enhanced wall treatment has been used for turbulence modeling.

Transport equations for k−e realizable turbulence model:

( )( ) [( ) ]t

j k b M kj j k j

k kku G G Y S

t x x x

µρ ρ µ ρεσ

∂ ∂ ∂ ∂+ = + + + − − +∂ ∂ ∂ ∂

and 2

1 2 1 3

( )( ) [( ) ]t

j bj j j

u C S C C C G St x x x kk

ε ε ε εε

µρε ε ε ερε µ ρ ρσ νε

∂ ∂ ∂ ∂+ = + + − + +∂ ∂ ∂ ∂ +

where

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1 max [0.43, ], , 2

5 ij ij

kC S S S S

η ηη ε

= = =+

Note that ε

is affected by the buoyancy is determined by the constant 3C ε

In these equations, kG represents the generation of turbulence kinetic energy due to mean velocity gradients,

bG

is the generation of turbulence kinetic energy due to buoyancy, MY

represents the contribution of fluctuation

dilatation in the compressible turbulence to the overall dissipation rate, 1C

and 2C are constants, kσ

and εσ are

turbulent Prandtl numbers for k

and ε respectively, kS and Sε are user- defined source terms. tµ

is defined as 2

t

kCµµ ρ

ε=

The difference between realizable k−e and the standard RNG k−e models is that Cµ is no longer constant. It is computed from

*

0

1

s

CkU

A Aµ

ε

=+

using the relationships *

ij ijij ijU S S= + Ω Ω , 2ij ij ijk kε ωΩ = Ω − , ijij ij kε ωΩ = Ω − , 1

2ji

ijj i

uu

x x

∂∂Ω = − ∂ ∂

where ijΩ is the mean rate-of-rotation tensor viewed in a rotating reference frame with the angular velocity kω .

The model constants 0A and sA are given by 0 4.04, 6 cossA A φ= = , where 11cos ( 6 cos )

3Wφ −=

using the relations 3

ij jk kiS S SW

S= , 0.5

i j

ijj i

u uS

x x

∂ ∂= + ∂ ∂ , and ij ijS S S=

The model constants used in this analysis are 1 21 .4 4 , 1 .9 , 1 .0 , 1 .2kC Cε εσ σ= = = = .

Table 1. Description of boundary conditions

Inlet boundary conditions: i. Type of boundary Velocity-inlet. ux = 40 m s−1 , uy = uz = 0

ii. Reynolds number 1.5×105

iii. Turbulence intensity 5% iv. Turbulent viscosity ratio 5

Outlet boundary conditions: i. Type of boundary Pressure-outlet

ii. Pressure-specified 0 Gauge pressure iii. Backflow turbulence intensity 5% iv. Backflow turbulent viscosity ratio 5

Walls boundary conditions: i. Type of boundary Smooth walls

ii. Shear condition No-slip Working fluid properties:

i. Working fluid Air ii. Density of working fluid 1.225 kg m−3

Dynamic viscosity of working fluid 1.602×10-4 N s m−1

4.3 Grid generation The discretization procedure for the 3-D solution domain has been performed with the help of an inbuilt solid

modeling and meshing tool with FLUENT 6.2, called GAMBIT. It employs body-fitted Non-orthogonal grid system

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which is shown in Fig. 3. Hexahedral meshing was done along with the growth factor in the normal direction to the wall. The growth of the mesh size helps to get finer mesh near the wall as well as coarser mesh away from the wall. The number of elements used for all the ducts was of the order of 150,000. Wall function approach (Spalding and Launder [19]) has been adopted for the near-wall treatment as it is computationally economical for high Reynolds number flows, robust and reasonably accurate.

4.4 Post processing The governing equations have been integrated over the control volume and discretized by second order

upwind scheme. Also the governing equations have been linearized implicitly. Pressure-velocity coupling has been done using the SIMPLE algorithm. A point implicit (Gauss–Seidel) linear equation solver has been used in conjunction with an algebraic multigrid (AMG) method to solve the resultant scalar system of equations for the dependent variable in each cell. The residuals and mass weighted average velocity at exit plane have been monitored

to check the convergence of the solution. The convergence criteria of 61.0 10−× for the residuals were set at. The

equations were solved with the unsteady solver with a time step which was typically 41.0 10−× .The computations

have been carried out in a Windows based Pentium-4 workstation with 512 Mb RAM. It took on an average 60-65 hours for each solve.

V. Results and Discussion For aircraft intake applications, the measure of duct aerodynamic performance is the ability to decelerate the

flow to the desired velocity, while maintaining high-pressure recovery and flow uniformity. Reduced pressure recovery lowers propulsion efficiency, whereas; non-uniform flow conditions at the engine face lower engine stall and surge limits. Detailed flow measurement within this diffuser has been carried out. It shows a non-uniform velocity distribution at its exit.

In order to uniform the flow pattern at diffuser exit, a two bladed ‘flow-deflector’ has been installed at the inlet of the diffuser. A flow visualization study by single point silk tufts (Paul et.al. [13]) revealed that the best flow pattern was available at 90° orientation of the deflector (measured anti-clockwise from horizontal plane). At this orientation, the counter-rotating vortices are not prominent, since the steady and unsteady turbulent flow region increases, especially near the wall of the diffuser. This is due to the installation of deflector, which ‘kills’ the pattern of counter-rotating vortices and makes an improvement of the flow at diffuser exit.

Major flow quantities, like mean velocity, radial components of mean velocity, total pressure and, static pressure at the exit of the diffuser have been measured intensively with the help of a pre-calibrated five-hole pressure probe and presented here in contour forms. Comparison of flow distribution at exit, with or without using flow deflector has also been carried out both experimentally and computationally. But for the convenience, experimental results at the diffuser exit using flow deflector have been presented here and compared with the computational results

5.1 Mean velocity contours The normalized mean velocity contours at the exit of the test sections of the diffuser has been shown in Fig. 4.

Due to centerline curvature, a radial imbalance of the centrifugal pressure force ( )2 / cmU R is set up between the

wall-A (concave-CC) and wall-B (convex-CV), and the acceleration produced ( )2 / cU R acts radially inwards to the

duct. Hence, a pressure gradient is set up between these two walls and is responsible for shifting of high velocity fluid from wall-A to wall-B. Mass of flow has shifted from wall-A to wall-B, and hence, low velocity accumulates near wall-A.

The mean flow-field in the diffusing curved bend is dominated by a pair of counter-rotating streamwise vortices (Lin et.al. [2]) set up by the cross-stream pressure gradient in the bend, which balances the centrifugal force of the fluid as it is diffused and turned. The low velocity fluid in the boundary layers on the wall-B is forced towards the wall-A along the circular wall by this pressure gradient. Flow continuity, then requires a high velocity fluid in the central portion of the duct exit flowing from wall-A towards wall-B. The modest diffusion present here is manifested as a general deceleration of this flow portion. The works reported by Whitelaw et.al. [20] supports the above discussion.

In Fig. 4 (b), the exit cross-section mostly covers with 0.4 U mean velocity, whereas, near to wall, the velocity varies between 0.2U and 0.3U . However, in the case of bare duct [Fig. 4 (a)], the low velocity zone (0.3U ) concentrates towards the wall-A; and hence, no uniform velocity profile is seen across the section. A comparison of

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Fig. 4 (a) and (b) reveals that a higher velocity zone (i.e. 0.7 U ) at the exit of the bare duct shifts from wall-B to wall-A. It can be attributed to the effect of the swirl angle of the blades of deflector, which somewhat shifts the flow from wall-B to wall-A. Fig. 4 (b) also depicts that the flow is more uniformly distributed at the exit using flow deflector at the inlet, and the same is also supported computationally in Fig. 4 (c).

5.2 Total pressure contours The normalized total pressure contours at the exit plane for bare duct as well as duct with flow deflector are

shown in Fig. 5 (a) and (b) respectively. It is evident from the Fig. 5 (b) that with deflector, the higher total pressure fluid shifts from wall-A to wall-B. This is attributed to the effect of deflector blade (especially the swirl angle of the deflector), which rotates the flow to some extent. The overall flow pattern at the exit of Fig. 5 (b) is rotated from wall-B to wall-A. However, the high pressure is recorded in case of Fig. 5 (b) with deflector is only 0.2 dp as

compared to 0.5 dp observed in the bare duct [Fig. 5 (a)]. This is due to the blockage of inlet mass flow by the deflector blades. However, a better uniform flow of total pressure has been achieved at the exit with installing flow deflector as compared to the bare duct. Similar uniform nature is shown in the case of mean velocity contour in Fig. 4 (b) as discussed earlier.

A weak pressure zone is also noticed in a confined space near the bottom of the wall as shown in Fig. 5 (b). This is the region, where flow is suspected to behave in some irregular way .This is probably due to from the naturally occurring vortices, convecting low momentum fluid away from this side of the duct wall. Reichert & Wendt [5] also observed similar nature.

5.3 Cross flow velocity plot The cross-flow velocity vectors of 30°/30° diffuser with the flow deflector have been plotted in Fig 6. The

cross-flow rotates in the upper-half of the wall-A in a counter-clockwise direction, whereas another vortices is formed in the lower half of the wall-B and rotates in a clockwise direction. The overall direction of the cross-flow pattern is from wall-A to wall-B. However, in some regions near the bottom wall, the flow reverses, which may be the indicator of some ‘disturbed’ flow-field. It can be said that the secondary flow generated by the flow deflector overwhelms the natural secondary flow caused in earlier cases of bare duct due to the centerline curvature.

The cross-flow distortion is largely improved with the installation of flow deflector that destroyed the pair of vortices (caused by the naturally occurred secondary flow due to centerline curvature) as evident from the flow-visualization studies carried out by Paul et.al. [5].

5.4 Wall Static Pressure Distribution To have an idea of static pressure variation along the walls; pressure taps of two opposite curved sides (wall-

A and wall-B) on the plane of curvature of the 30°/30° diffuser have been used. Readings have been taken with the help of a multi-channel electronic pressure scanner by changing the orientations of the flow-deflector (viz. 0°, +45°, +90°, −45° and bare duct), and the results have been plotted in a non-dimensional form as shown in Fig. 4 & 5. The wall static pressures have been normalized by the inlet dynamic pressure (i.e. 20.5 Uρ = 896 N/m2). During experiments, the inlet Reynolds number for all cases have been fixed at 1.5×105, which corresponds to the air velocity of 40 m/s. It is also to be noted that the curves on wall-A are smooth compared to wall-B. This may be due to accumulation of low velocity fluid at the point of inflexion ( iL D =3.0). Similar variation was also reported by Majumdar et.al. [6]. Hence, the two predominant features of the pressure distribution are observed on the two walls of the diffuser are:

• An overall increase in static pressure on the walls with stream-wise position of the mean pressure level due to gradual increase in area.

• The cross-stream gradient in reaction to the centrifugal force of the flow, as it is turned. The balance of these effects result in a continually adverse pressure gradient up to iL D = 2.5 to 3.5 on the CV portion of the wall-A followed by a mild acceleration as the flow relaxes in the CC portion and through the tailpipe. Whereas, on the CC portion of the wall-B, there is an initial region of acceleration up to iL D =

0, which is then followed by a continuing adverse pressure gradient up to the inflexion point ( iL D = 3.0).

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VI. Conclusions

From the present investigation, the following conclusions have been drawn: • Flow visualization study suggests that the best orientation of the flow deflector at which the exit flow

becomes uniform, lies between 90o to −45o measured counter-clockwise horizontally. At this position, no distorted flow in the form of separation or fixed stall is observed. This signifies that the separation could be reduced largely with the installation of flow deflector.

• For the bare duct, (i.e., diffuser without flow deflector), a secondary motion in the form of counter-rotating vortices is generated within 30°/30° diffuser. But with the installation of the flow deflector at the upstream of the diffuser inlet, the counter-rotating vortices at the exit plane has been destroyed and the flow pattern has been re-oriented.

• The bulk-shifting of the flow from wall-A to wall-B is very distinct at the exit plane of the diffuser. • Flow pattern at the exit appeared more uniform with the installation of flow deflector as revealed from

the figures of mean velocity contour (Fig. 2). • A more fluctuation of wall static pressure is noted on wall-B than that of wall-A because of

accumulation of low velocity fluid at the point of inflexion ( iL D = 3).

References

1. R.W. Guo, J. Seddon, Swirl characteristics of an S-shaped air intake with both horizontal and vertical offset, Aeronautical Quarterly, (May 1983) 130-146. 2. Qi Lin, R.W. Guo, Vortex control investigation of swirl in S-shaped diffuser, Astronautica Sinica, 10 (1) (1989). 3. P.F. Weng, R.W. Guo, A new approach to swirl control in an S-duct, Proc. of the Seminar on Aero-engines, Nanging Aeronautical Institute, P.R. China, (April 1991). 4. P.F. Weng, R.W. Guo, New method of swirl control in a diffusing S-duct, J. of AIAA, 30 (7) (1992) 1918-1919. 5. B.A. Reichert, B.J. Wendt, Improving curved subsonic diffuser performance with vortex generators, J. of AIAA, 34 (1) (1996) 65-72. 6. B.J. Wendt, B.A. Reichert, Vortex ingestion in a diffusing S-duct inlet, J. of Aircraft, 33 (1) (1996) 149-154. 7. B.H. Anderson, J. Gibb, J, Vortex generator installation studies on steady state & dynamic distortion, J. of Aircraft, 35 (4) (1998) 513-520. 8. R.G. Dominy, D.A. Kirkham, A.D. Smith, Flow development through inter-turbine diffusers, Trans. ASME, J. of Turbo- machinery, 120 (1998) 298-304. 9. B.H. Anderson, A robust design methodology for optimal micro-scale secondary flow control in compact inlet diffusers, 2002, AIAA 2002-0541. 10. A.M. Pradeep, R.K. Sullerey, Secondary flow control in a circular S-duct diffuser using vortex generator jets, AIAA 2nd

Flow Control Conference, Portland, Oregon, 2004, AIAA Paper 2004-2615. 11 . R.K. Sullerey, V. Mangat, A. Padhi, Flow control in serpentine inlet using vortex generator jets, AIAA 3rd Flow Control Conference, San Francisco, California, 2006, AIAA Paper 2006-3499. 12. M.L. Mathur, A new design of vanes for swirl generation, Institution of Engineers (India) J., M.E, 55 (Nov. 1974). 13. A.R. Paul, Dr. B. Majumdar, Studies on flow deflector installation at inlet of sigmoid ducts, Proc. of 2nd International Conf. on Fluid Mechanics & Fluid Power, Indian Institute of Tech., Roorkee, India, Vol. I, Dec. 13-15, 2002, pp. 209-216. 14. R.W. Fox, S.J. Kline, Flow regimes in curved subsonic diffusers, Trans. ASME, J. of Basic Engineering, 84 (1962). 15. D.W. Bryer, R.C. Pankhurst, Pressure-probe methods for determining wind speed and flow direction, National Physical Laboratory, London, 1971. 16. FLUENT, User’s Guide, CFD Software Package, Ver. 6.2, Fluent Inc, Lebanon, NH, 2005. 17. G. Biswas, The k−e

model, the RNG k−e

model and the phase-averaged model, in: G. Biswas and V. Eswaran (Eds.), Turbulent Flows: Fundamentals, Experiments and Modeling, IIT Kanpur Series of Advanced Texts, Narosa Publishing House, India, 2002. 18. D. Ravi Kumar, S. Bandyopadhyay, Computation of three dimensional compressible turbulent flow through a non- diffusing S-duct, 5th Annual CFD Symposium, Aeronautical Society of India, Bangalore, 9-10 August, 2002. 19. B.E. Launder, D.B. Spalding, The numerical computation of turbulent flows, Computer Methods in Applied Mechanics and Engineering 3 (1974) 269-289. 20. J.H. Whitelaw, S.C.M. Yu, Turbulent flow characteristics in an S-shaped diffusing duct, Flow Measurements & Instrumentation, 3 (3) (1993) 171- 179. 21. A.D. Vakili, J.M. Wu, P. Liver, M.K. Bhat, Measurement of compressible secondary flow in a circular S-duct, AIAA paper: 83- 173, July 1983. 22. S.R. Wellborn, B.A. Reichert, T.H. Okiishi, An experimental investigation of the flow in a diffusing S-duct, AIAA paper: 92- 3622, July 1992.

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Fig. 1. Twin bladed flow deflector.

Fig. 2. Schematic of 30o/30o sigmoid diffuser.

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Fig. 3. Numerical mesh for 30o/30o sigmoid diffuser with a deflector installed at the inlet.

Fig. 5. Total pressure contours at exit plane of 30o/30o diffuser.

I: Wall A, II: Wall B For a bare duct (Experimental). b) For a duct with deflector positioned at 90o (Experimental)

c) For a duct with deflector positioned at 90o (Computational)

Fig. 4. Mean velocity contours at the exit plane of 30o/30o diffuser.

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Fig .7. Static pressure distribution along wall-A

Fig .8. Static pressure distribution along wall-B

Fig. 6. Cross flow vector plot at exit plane of 30o/30o

diffuser.