EFFECT OF A NEW VORTEX GENERATOR ON THE … · BASELINE COMPRESSOR CASCADE In the present work, a...

Proceedings of the ASME 2015 International Mechanical Engineering Congress & Exposition

IMECE2015

November 13-19, 2015, Houston, Texas, USA

IMECE2015-52293

EFFECT OF A NEW VORTEX GENERATOR ON THE PERFORMANCE OF AN

AXIAL COMPRESSOR CASCADE AT DESIGN AND OFF-DESIGN CONDITIONS

Ahmed M. Diaa Assiut University

Assiut, Egypt [email protected]

Mohammed F. El- Dosoky Assiut University


Mahmoud A. Ahmed Assiut University


Omar E. Abdelhafez Assiut University


ABSTRACT Secondary flows are noxious to axial compressor

performance. To overcome and control those secondary flows,

vortex generators are used as a passive control device.

Controlling secondary flows will lead to a further improvements

in the compressor performance. A new design of vortex

generator is considered in this investigation in order to control

secondary flows in axial compressor cascade at design and off-

design conditions. Numerical simulations of a three-

dimensional compressible turbulent flow have been performed

to explore the effect of the vortex generators on the reduction of

secondary flows. Six different incidence angles are used for the

off-design operation investigations. Based on the simulation

results, the pressure, velocity, and streamline are used to follow

up the development of the secondary flows. Thence, total

pressure loss coefficient, static pressure rise coefficient,

difference in flow deflection angle, and diffusion factor are

estimated. Results indicate that vortex generators have a

significant effect on the development of secondary flows at off

design operation as they cause a reduction in total pressure loss,

they also affect the loading behavior of the cascade as they

cause a slight change in the cascade deflection, and a slight

decrease in the diffusion factor which causes unloading of the

blade. Static pressure rise is significantly reduced at negative

incidence angles while a slight reduction occurs at positive

incidence angles. In a word, the new design of the vortex

generator enhances the cascade aerodynamic performance and

enlarges the operating range of the cascade towards the positive

incidence region.

INTRODUCTION Axial compressor gains importance because of its

relevance to gas turbine applications in military and commercial

airplanes and its usage in stationary power generation.

Therefore, early researches have been carried out to enhance its

overall performance. The secondary flows have noxious effects

on the axial flow compressor due to extract energy from the

working fluid and increase the flow instability. Therefore,

controlling secondary flows will extremely improve the

aerodynamic performance of the compressor. The components

of the secondary flow in the axial compressor are the endwall

boundary layer separation, the horseshoe vortex, the corner

vortex, the tip vortex, the endwall crossflow, and the passage

vortex. Since the early 1950s, numerous researches were carried

out [1–9] to investigate secondary flow phenomena. These

investigations played a decisive role in improving the

compressor performance. Recently, many researchers

investigated the impact of the three-dimensional blades and

their endwall boundary layer separation as well as flow

separation in the corners of the blade passages on the

development of the secondary flows [10-12]. In pursuit of

controlling secondary flows, both passive and active methods

have been utilized to reduce or overcome the effects of

secondary flows in axial compressors. Passive control methods

remain the preferable as simple and inexpensive techniques

[13,14]. Different configurations of passive flow control

devices were used and investigated such as slotted blading in

linear cascades [15], vane and plow vortex generators placed on

several positions [16], counter rotating and co-rotating

rectangular, triangular, and parabolic vane type vortex

generators [17–19], cavity as a control of shock wave

interactions with the turbulent boundary layer [20], low profile

vortex generators to reduce the boundary layer thickness [21],

and doublet vortex generators [22]. An extensive review of

boundary-layer flow-separation control by micro vortex

generators (low profile vortex generators) had been compiled as

reported in references [23-24]. Low profile vortex generators

are used to energize the low momentum flow near the wall

surface without energy expenditure during the momentum

transfer from the free-stream flow to the near wall region. Yet,

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2 Copyright © 2015 by ASME

this leads to an overwhelmed stronger adverse pressure gradient

and therefore avoids or delays the flow separation. In case of

turbulent flow over a flat plate, experimental results indicated

that the vane and wheeler type of vortex generators can

efficiently reduce the flow separation. Using the vortex

generator height (h/δ) of 0.1 to 0.4 was fairly efficient with

much reduction in the drag effect [25]. Furthermore, vane type

with height (h/δ) of 0.8 achieved the largest pressure recovery

[26]. An experimental comparison between two passive

methods for controlling shock induced separation on a turbulent

flat plate boundary layer was carried out by McCormick [22]. A

doublet wedge type vortex generator with h/δ=0.36 was used

versus passive cavity (porous wall with a shallow Cavity

underneath). It is reported that the low profile vortex generators

were found to be significantly suppressing the shock induced

separation and improve the boundary layer characteristics

downstream the shock whereas the mass-averaged total pressure

loss reduction increased.

Recently, experimental and numerical studies of the

effect of two vortex generator types with different

configurations on the performance of the compressor cascade

were conducted by Hergt et al. [27]. They concluded that using

different types of vortex generators results in significant

reduction in cascade losses and loading, which in turn shifts the

compressor operating range towards higher positive incidence

angle. They also achieved a total pressure loss reduction by 9%

at design operation. Moreover, vortex generators have a

significant effect on the cascade deflection and a remarkable

enhancement of the cascade stall range. However the static

pressure rise due to inserting a vortex generator was nearly

unaffected. Diaa et al [28-29] have investigated the effect of the

curved side vortex generators as a control device with axial

compressor cascade, they reported a reduction of total pressure

loss of 12%.

Based on the literature survey of the research work related to

the current subject, the suggested curved surface design of

vortex generator is inspired from Wheeler's doublet VG [30]

that was used in several applications on flat plates. It was found

that using doublet VG significantly enhances the mixing process

between the low and high momentum streams. In addition, the

total pressure recovery was found to be increased, and the onset

of separation was delayed or eliminated. Further modification

on the curved side VG design is added by including rounded

nose rather than sharp nose. The new design achieved the

highest total pressure recovery of about 20.7%. More details

can be found in the article.

NUMERICAL METHODOLOGY

BASELINE COMPRESSOR CASCADE In the present work, a linear high speed compressor

cascade that was reported by the research group of

Hergt et al. [24] is used. Their compressor cascade was

designed by “MTU Aero Engines” with the profile shown in

fig.1. The design parameters and the operating conditions of the

cascade are summarized in Table 1.

FIGURE 1 COMPRESSOR BLADE PROFILE.

TABLE 1 DESIGN PARAMETERS AND OPERATING

CONDITIONS

Mach number at inlet M1 = 0.66

Inlet flow angle β1=132°

Turning angle Δβ = 38°

Stagger Angle βst =105.2°

Blade chord C = 40 mm

Blade span L = 40 mm

Pitch to chord ratio s/c = 0.55

Endwall boundary layer thickness δ = 4 mm

Computational domain Computational domain and its boundaries are shown in fig.2.

The non-slip boundary condition is applied at the walls

representing the top boundary, the bottom boundary (Endwall),

and the blade surfaces demonstrating the suction and pressure

sides including the leading and trailing edges. Periodic

boundary conditions are applied on the domain sides. The

pressure outlet boundary condition is defined at the outlet plane.

Figure 2 Boundary conditions of the computational domain

(a top view showing inlet, outlet, blade suction and pressure surfaces, and periodic boundaries.

Fully developed flow is adopted at the inlet with an average

Mach number of 0.66, while the inlet flow angle is varied as

will be shown later.


Numerical solution The investigated numerical cases present the effect of the

vortex generators on the cascade performance at design and off-

design conditions. Therefore, six different sets of vortex

generators are investigated at six different incidence angles (i)

ranging from i=+6⁰ to i= - 6⁰ with step of 2⁰. Commercial

solver ANSYS FLUENT is used as a CFD tool to solve steady

Reynolds-averaged Navier Stokes fully coupled with transition

shear stress transformation (SST) turbulence model. The

transition SST model is used in this simulation because of its

ability to accurately capture the transition and separation flows.

Three-dimensional grid is constructed using structured mesh of

H-O-H topology as shown in fig.3.

FIGURE 3 THREE DIMENSIONAL MULTI-BLOCK MESH

WITH ENLARGED TOP VIEW OF THE MESHING AROUND THE BLADE AND TWO MAGNIFIED VIEWS OF MESHING

AROUND LEADING AND TRAILING EDGES.

In the present study several criteria are used to measure the

grid convergence. The first one is the grid independency tests

which are shown in Figs. 4, and 5.

Solution with grid independency is reached by testing four

different meh sizes ranging from 0.4 to 1.6 million cells. As

shown in fig.4. Solution with grid independency is reached at a

mesh of 0.8 million cells.

Figure 5 shows the grid convergence test using the velocity

contours at the exit plane. This test is carried out by using

different grid size starting with a coarse grid with 0.4 Million

cells up to a fine grid with 1.6 million cells. It is shown that the

values of velocity contours is not subjected to a significant

change beyond the 0.8 Million cells grid, and that is another reason to adopt this grid in the present simulation.

The second criteria is performed based on the value of

residual of different quantities to be less than or equal to 5.0E-

5. The last one is based on calculations of Grid Convergence

Index (GCI) as reported by Roache[31]. The GCI measures the

percentage of the deviation between the computed values and

the asymptotic values. It specifies the error band and how far

the solution is from the asymptotic value. Also, it shows how

much the solution would change with grid refinement. A small

value of GCI indicates that the computation is within the

asymptotic range. Therefore, Three different meshes with

195000, 420000, 826000 are selected to determine GCI. Table

2 shows the GCI values for axial velocity and static pressure.

For the axial velocity, the GCI is found to be 0.53% for the

highest grid size, in addition for the static pressure the CGI

value is 0.51% for the finest grid. Table 2 shows the calculation

of the GCI for different mesh size. As listed in Table 2,

successive grid refinement results in a reduction in GCI.

Therefore, the solution of the finest grid (826,000) cells is grid

independent.

Table 2: grid convergence index for static pressure and axial

velocity

GCI21 GCI32

Φ=static pressure 6.4% 0.51%

Φ=axial velocity 3.78% 0.53%

Figure 4 Mass averaged velocity at exit plane (located 40%

of chord length behind trailing edge) of different mesh size starting from 0.4 to 1.8 million cells.

Figure 5 velocity contour lines (m/s) at exit plane for

different grid sizes (0.4 – 1.6 Million cells).


Therefore, the present simulation is performed using 0.8

million cells to get results that independent of the grid size and

to reduce the computational time. The minimum cell height

near the walls is adopted to have y+<1, which is considered to

capture and resolve the boundary layer at the blade surfaces and

enwalls.

To conclude, a new design of vortex generator (VG) is

inserted into the flow passage of a compressor blade cascade,

and its effect on the compressor cascade performance is

investigated for both design and off-design conditions. For the

off design conditions incidence angle will be varied from i= - 6⁰ to i= +6⁰ with step 2⁰.

The new VG is shown in fig.6. Two main configurations of

the new VG are investigated. First one is having a sharp edged

nose as shown in five stets A, B, C, D, and E with different

geometrical dimensions. The second configuration is set F

which has a rounded edge nose. Four different ratios of r/δ=

0.25, 0.5, 0.75 and 1.0 are investigated, but the only reported

results in this manuscript is the case with r/δ=0.5; since it

achieves the highest performance relative to the others. The six

different sets with different geometrical dimension ratios are

summarized in Table 3.

FIGURE 6 TWO CONFIGURATIONS OF THE NEW

VORTEX GENERATOR, (A) WITH SHARPED EDGE NOSE AND (B) WITH ROUNDED EDGE NOSE.

Table 3 Summary of vortex generators geometrical ratios

represented in set A, B, C, D, E and F.

Set h/δ w/h e/h r/δ

A

0.4

3

6 __ B 4

C 5

D

6 E 5

F 0.5

VALIDATION In order to validate the current results, a comparison with

Hergt's [27] results is performed. Isentropic Mach number

distribution on the blade profile at midspan is compared with

isentropic Mach number calculated from the experimental work

by Hergt [27] as shown in fig. 7. The Figure shows a good

qualitative and quantitative agreement with Hergt's results.

FIGURE 7 COMPARISON BETWEEN THE MEASURED

ISENTROPIC MACH NUMBER DISTRIBUTION [27] AND THE PRESENT SIMULATION.

Another validation comparison is carried out using the

mass-flow averaged spanwise pressure loss distribution of the

base case as shown in fig. 8. The figure shows a good

qualitative agreement with maximum deviation of 7% near the

bottom end wall.

FIGURE 8 COMPARISON BETWEEN THE TOTAL PRESSURE

LOSS DISTRIBUTION AT A PLANE LOCATED AT 40% OF CHORD LENGTH BEHIND THE TRAILING EDGE [27], AND

THE PRESENT SIMULATION.

RESULTS AND DISCUSSION In this section, the numerical results with vortex generators

at both design and off-design conditions are presented and

discussed.

In particular, The study focuses on the streamlines at blade

suction surface and other performance parameters such as, total

pressure loss coefficient, flow deflection angle, static pressure


rise coefficient, and diffusion factor which will clarify the effect

of VG on the cascade performance.

Effect of vortex generator on secondary flows In order to determine the separation lines, secondary flow

components such as the footprint of the propagation of the

passage, corner, and horseshoe vortices are considered. The

limiting streamlines on the blade suction surface are used to

trace the changes in the separation lines after using vortex

generator. In Fig.9, streamlines at blade suction surface with

and without vortex generators show the position of both the

separation lines, the attachment lines and the recirculation

volumes .

For base case without vortex generator, the passage vortex

is firstly separated from the endwall and then reattached to the

suction surface at AL1. The separation line SL1 indicates the

separation of the passage vortex from the suction surface

followed with reverse flow over the blade which is originated

near the midspan towards SL1, and then the flow is separated

from the blade surface by the passage vortex as mentioned

before. In addition, separation line SL2 indicates the separation

of the corner vortex along the blade suction surface, while AL2

shows the reattachment of the corner vortex with the blade

surface. For cases, where VGs are used, the vortex structure

inside the flow passage is affected by the two generated counter

rotating vortices.

For case A, SL1, and AL1 are both moved toward the

trailing edge which is an indicator for the deflection of the

passage vortex into the flow direction. This deflection is due to

one of the generated vortices near the suction side, also, SL2

and AL2 disappeared from the streamlines of this case. Reverse

flow on the blade surface is decreased for case A, as shown in

Fig. 8. As w/h increased for design B, no significant changes in

the position and distribution of the streamlines are identifiable.

Further increase in w/h, as in case C, causes a larger

deflection of the passage vortex which is indicted by the

movement of SL1 and AL1 downstream towards the trailing

edge. Reverse flow on the suction side is diminished.

Another increase in w/h for case D yields a disappearance

of SL1 and AL1, furthermore, the reverse flow also

disappeared, these changes due to increasing w/h are attributed

to a more strength generated vortex by inserting the VG.. In

addition, the movement of the separation lines also indicates to

an increase in stretching rate of the vortex control region with

higher w/h values. Streamlines behavior of case E is similar to

that caused by case F; therefore it is not reported here.

In case F with the rounded nose edge, separation and

attachment lines SL1 and AL1 are vanished. In addition, reverse

flow is noted on the blade surface.

On the top half of the blade surface, for base case a

circulation bubble CB is noticed in the upper and lower half

of the blade span. After using VG, a circulation bubble CB1 is

noticed on the upper half of the blade span suction surface. This

circulation bubble is formed by the interaction of the cross

flow driven by the adverse pressure gradient at the top wall

corners and midspan and the streamwise flow.

FIGURE 9 SKIN FRICTION LINES ON THE SUCTION

SURFACE FOR BASE CASE (WITHOUT VG) AND CASE A, B, C, D, AND F AT DESIGN OPERATION ( i=0 [deg.]).

Effect of vortex generator on total pressure loss

coefficient Total pressure loss coefficient (TPLC) is known as a

compressor performance parameter which is used to judge the

loss behavior of the flow. TPLC is calculated from the

following equation:

11

21 ),( ),(

st

tt

pp

zyppzyTPLC

(2)

Figure 10 shows local TPLC contours at exit plane (which

is located 40% of chord length downstream the trailing edge) at

design and off design. From this TPLC contours, it is clear that

at high negative and positive incidence angles (when i=-6 and

+6[deg.]) TPLC is increased when using VG in comparison

with the base cascade. By decreasing incidence angle to i=-4


and +4[deg.] a slight decrease in TPLC is noted in all cases

with VG compared with the base case. Further reduction in

TPLC is gained by decreasing the incidence angle to i=-2 and

+2[deg.].Therefore using VG will decrease the total pressure

loss over a range of incidence angles between i=-4⁰ to +4⁰. Total pressure loss factor (TPLF) represents the difference

of total pressure loss coefficient (between case with VG and the

base case) normalized by the total pressure loss coefficient of

the base case as stated in equation 3. To calculate the TPLF, the

local value of TPLC(y,z) in equation 2 is integrated in the

spanwise direction and pitch wise directions for base and all

VG cases.

Base

BaseVG

TPLC

TPLCTPLCT

PLF (3)

Figure 10 TPLC contours measured at exit plane (located

at 40% of chord length behind trailing edge) at incidence angles i=-6, +4,-2, +2, +4, and +6 [deg.] for Base, D, E and F

cases.

Figure 11 shows the TPLF for all cases. At design

condition, for cases A to D, where β1=132 deg., it is noticeable

that by increasing w/h the TPLF is reduced which indicates a

reduction in the total pressure loss in comparison with the base

case. This is due to the increase of the stretching rate of the

generated vortices. The TPLF ranges from -5.3% to -12.6 % for

case A and D respectively. For case E a further reduction in

TPLF is attained with value of -18.6 %.The highest reduction is

reached using case F with rounded nose edge, where the TPLF

value is- 20.7%.

Figure 11 Total pressure loss factor with different incidence

angles for set A, B, C, D, E, and F[%]

At off-design condition, it is clear that TPLF increases with

the increase of incidence angle, for negative incidence angle

i=-6 deg. (β1=126 deg.) there is no reduction in total pressure

loss attained in all cases, instead an increase in the total

pressure loss is achieved by using vortex generator relative to

the base case. At negative incidence angles, the two generated

vortices will not have the same strength as at design condition.

The pressure side generated vortex is strong, but the suction

side generated vortex is far too weak. This non-symmetric

strength of the two generated vortices will cause an opposed

effect on losses reduction, since the weaker vortex will not be

able to mix low and high momentum flows properly. By

decreasing the incidence angle to i= - 4 where β1=128 deg., the

non-symmetric strength still exists but with smaller difference

between the generated vortices, and this is the reason why TPLF

is slightly reduced at this operating point. Furthermore at i= -2

β1=130 deg. TPLF is further reduced for all cases, the highest

reduction is reached at case E with value of -12.3 %.

At positive incidence angles the effect is reversed where

the suction side generated vortex is stronger than that of the

pressure side . TPLF is increased for all cases at i=+2 where

β1=134 deg., and the uppermost value of TPLF reduction at this

incidence angle occurs at case E with value of -12.8 %. TPLF is

further increased with i=+4 where β1=136 deg., for all cases. At

i=+6 deg., where β1=138, all cases have an increase in pressure

loss which is noted by the positive vale of TPLF for all cases at

this incidence angle. Therefore, at design conditions, case F

causes the highest reduction in TPLF; nevertheless, at off-

design operation case E yields higher reduction in TPLF than

this caused by case F. Consequently, optimization of the best

case to be used to cover a wider range of operation with the

highest reduction in pressure losses is still needed.


Effect of vortex generator on flow turning angle Flow turning angle is computed from Equation 4. In order

to track the change in flow deflection by using vortex

generators, the difference of the flow turning angle will be

normalized by flow turning of the base case as shown in

Equation 5.

21 (4)

Base

BaseVG

__

(5)

The normalized difference of flow turning angle is shown

at midspan for base case and all VG cases in Fig. 12. At design

operation, the flow turning angle is either decreased or

increased by using VG where the maximum reduction is of

- 0.25 % with the case C and the maximum increase of 0.38%

with the case D. Consequently, a slight change is noticed in the

flow turning angle for all cases at this point of operation.

At off-design condition, a high negative incidence angle

i= - 6 deg., causes an increase in the flow turning angle with

value of 1.8 % at case A, slight change in this value is noted for

other cases. As negative incidence decreases to i= -4 deg., flow

turning is increased to 1.2% at case B. Further decrease in

incidence angle to i= -2 deg. flow turning is lowered to 0.53 %

at case B. At positive incidence angles, the change in flow

turning angle is lower than that caused by negative incidence.

At i= +2 deg, a slight increase in flow turning angle is

identifiable with intermediate value of 0.32 % at case B. By

increasing the positive incidence to i= +4 deg, further increase

in flow turning is attained with middle value of 0.45% at case

A. Highest positive incidence angle of i=+6 deg results in an

increase flow turning with central value of 0.63% at case A

.Consequently, using VG has no significant effect on the flow

turning angle at design and off-design operation.

Effect of vortex generator on the diffusion factor Blade loading is assessed by the diffusion factor (DF)

which relates to the peak velocity on the suction surface of the

blade to the velocity at the trailing edge. The Diffusion factor

can be defined by Equation 6[33] .

11

2

21

DF (6)

Normalized diffusion factor in Equation 7 is used to

measure the effect of different cases with VG at different

incidence angles on the diffusion factor.

Base

BaseVG

DF

DFDFDF

___

(7)

Figure 12 Normalized difference of flow turning angle at

midspan for case A, B, C, D, E, and F with various incidence angles [%].

The Diffusion factor is used as an indicator of the

probability of stall, since the increase of DF leads to more

possible separation [34]. Figure14 shows variation in DF for all

cases at the entire operating range.

For sets A, B, C, D, and E , diffusion factor decreases over

the whole operating range and it almost remains constant except

for i=+6 , where DF increases. For set F, diffusion factor is

slightly increased for both positive and negative incidence

angles as shown in Fig.13.

The first two terms of the diffusion factor defined by Eq. 6

represent the deceleration of the flow, while the third term

represents the flow turning [35]. For positive incident angles,

the declaration of the outlet flow is increased for almost all sets

which increases the deceleration term in the DF equation. It is

also noted that flow turning term is slightly decreased at this

range of incident angles as shown in Fig.14. The result is a

slight increase in the DF at this range of operation. At negative

incident angles, flow is less decelerated which allows the

deceleration term in the DF equation to decrease. The flow

turning term is also highly increased at this region as shown in

Fig.14, which is also reflected in the increase of the DF in the

negative incident region.

Figure 13 Diffusion factor for case A, B, C, D, and E with

different incidence angles [%]


Figure 14 Deceleration term , flow turning , and diffusion factor DF with different incidence angles for set F.

Effect of vortex generator on static pressure rise Static pressure rise coefficient is calculated form

Equation 8 and the normalized difference is calculated as shown

in Equation 9.

11

12

st

ss

PP

PPCp

(8)

Base

BaseVG

Cp

CpCpCp

___

(9)

Figure 15 shows the difference in static pressure rise

coefficient for all cases. As noticed, at negative incidence

angles, the static pressure rise is highly decreased while at

positive incidence angles static pressure is slightly decreased.

Similar effect is noticed for set F as shown in Fig.15.

The rise of Cp at β1=134 is probably attributed to the

difference of the axial velocity density ratio (AVDR) between

both β1=134 and the other positive incidence angles of 136,

138. It was found that the value of AVDR at β1 of 134 is less

than that of 136, and 138. Accordingly, the flow contraction

at β1=134 is less than that of 136, and 138. This means that

the flow is more decelerated at 134 than that of 136, and 138.

Consequently the value of Cp at 134 is greater than that of

136, and 138. As reported in reference [4], the AVDR is

defined by Eq. 10 as follows:.

2 2

1 1

x

x

vAVDR

v

Table 4: AVDR values at different inlet angles

Angle Set F Base Change in AVDR

β1=134 1.09291 1.08583 0.6517%

β1=136 1.12137 1.10868 1.1446%

β1=138 1.13321 1.12218 0.9833%

(10)

Figure 15 Static pressure rise coefficient for set A, B, C, D,

E, and F with different incidence angles [%]

CONCLUSIONS Numerical simulations of a three-dimensional compressible

turbulent flow have been performed to explore the effect of the

vortex generators on the reduction of secondary flows. Six

different sets of a new design of vortex generator are used in the

current work as a passive boundary layer control device in an

axial compressor cascade. The investigation is carried out over

a wide operating range of incidence angles at design and off

design conditions. The inlet Mach number is of 0.66 with inlet

angle of β1=132⁰ at design operation. For off-design operation

the flow incidence angle is changed between i= -6 and i=+6

deg. Numerical simulation is performed using commercial

software ANSYS FLUENT. The results show that for all

investigated configurations, a significant reduction of total

pressure loss up to 20.7% by case F is attained in comparison

with the base cascade without vortex generator at the design

point. In addition, this reduction in total pressure loss occurs at

incidence angles of i=±2, ±4 deg with maximum value of

12.8 %, while at i = ±6 deg an increase in total pressure loss is

noted.

For other investigated performance parameters, the flow

turning angle has a slight change for the different operating

range angles. Also, a slight decrease in the diffusion factor is

noticed for the entire operating range when using vortex

generators. Static pressure rise coefficient is rapidly decreased

by increasing the negative incidence, while it remain constant at

positive incidence angles. Therefore, using this vortex generator

reduces the total pressure loss and enlarges the operating range

of compressor cascade towards the positive incidence angle

region without a significant effect on the flow turning angle and

diffusion factor.


NOMENCLATURE

Symbols AVDR :Axial velocity density ratio

c : Chord length

Cp : Static pressure coefficient

DF : Diffusion factor

e : Vortex generator's length

GCI :Grid Convergence Index

h : Vortex generator's height

i : Incidence angle

L : Span

M : Mach number

P : Pressure (total or static)

r Rounded nose radius

S : Pitch

TPLC : Total pressure loss coefficient

TPLF : Total pressure loss factor

w : Vortex generator's width

β : flow angle

δ : Boundary layer thickness

ν : Relative velocity

ρ :Density

σ : Solidity

Subscripts is : Isentropic

1 : At inlet

2 : At measurement plane

s : Static pressure

t : Total pressure

Base : Base case without vortex generator

VG : Case with vortex generator

θ : Tangent component

x : Axial component

REFERENCES

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Visulaization study of secondary flows in cascades,”

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[3] B. Lakshminarayana, and Horlock, J. H., 1963,

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287–307.

[4] H. Schlichting, and Das, A., 1966, “Recent Research

on Cascade-Flow Problems,” pp. 221–228.

[5] Gessner, B. F. B., 1972, “The origin of secondary flow

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[6] Wisler, D. C., 1985, “Loss Reduction in Axial-Flow

Compressors Through Low-Speed Model Testing,”

(3).

[7] Dong, Y., 1987, “Three-Dimensional Flows and Loss

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Lakshminarayana, B., 1966, “Wall Stall in Compressor

Cascades,” Journal of Fluids Engineering, 88(3), pp.

637–648.

[9] Greitzer, E. M., 1980, “Review—Axial Compressor

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EFFECT OF A NEW VORTEX GENERATOR ON THE … · BASELINE COMPRESSOR CASCADE In the present work, a...

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