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Comparative studies have been carried out on two axial flow fan rotors of controlled vortex design (CVD), at their design flowrate, in order to investigate the effects of circumferential forward skew on blade aerodynamics. The studies were based on computational fluid dynamics (CFD), validated on the basis of global performance and hot wire flow field measurements. The computations indicated that the forward-skewed blade tip modifies the rotor inlet condition along the entire span, due to its protrusion to the relative inlet flow field. This leads to the rearrangement of spanwise blade load distribution, increase of losses along the dominant part of span, and converts the prescribed spanwise blade circulation distribution towards a free vortex flow pattern. Due to the above, reduction in both total pressure rise and efficiency was established. By moderation of the radial outward flow on the suction side, being especially significant for non-free vortex blading, forward sweep was found to be particularly useful for potential reduction of near-tip loss in CVD rotors. Application of reliable CFD-based design systems was recommended for systematic consideration and control of both load-and loss-modifying effects due to non-radial blade stacking.
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Aerodynamic effects of forward blade skew in axial flow
rotors of controlled vortex design
,ARAKwedikha, Cs Horváth, M Balczó, M M Lohász, and T Régert
Department of Fluid Mechanics (DFM), Budapest University of Technology and Economics, Budapest, Hungary
The manuscript was received on 6 February 2007 and was accepted after revision for publication on 11 July 2007.
DOI: 10.1243/09576509JPE420
Abstract: Comparative studies have been carried out on two axial flow fan rotors of controlled
vortex design (CVD), at their design flowrate, in order to investigate the effects of circumferential
forward skew on blade aerodynamics. The studies were based on computational fluid dynamics
(CFD), validated on the basis of global performance and hot wire flow field measurements. The
computations indicated that the forward-skewed blade tip modifies the rotor inlet condition
along the entire span, due to its protrusion to the relative inlet flow field. This leads to the
rearrangement of spanwise blade load distribution, increase of losses along the dominant part of
span, and converts the prescribed spanwise blade circulation distribution towards a free vortex
flow pattern. Due to the above, reduction in both total pressure rise and efficiency was established.
By moderation of the radial outward flow on the suction side, being especially significant for non-
free vortex blading, forward sweep was found to be particularly useful for potential reduction of
near-tip loss in CVD rotors. Application of reliable CFD-based design systems was recommended
for systematic consideration and control of both load- and loss-modifying effects due to non-
radial blade stacking.
Keywords: axial flow turbomachinery, controlled vortex design, forward blade skew, forward
blade sweep, circumferential forward skew
Rotors of axial flow turbomachines are often of ‘con-
trolled vortex’ design (CVD) [1]. This means that
contrarily to the classic free vortex concept prescribing
spanwise constant design blade circulation, the cir-
culation – and thus, the Euler work – increases along
the dominant part of the blade span in a prescribed
manner. CVD guarantees a better utilization of blade
sections at higher radii, i.e. it improves their contribu-
tion to the rotor performance. By this means, rotors
of high specific performance can be realized, i.e. rel-
atively high flow rate and total pressure rise can be
obtained even with moderate diameter, blade count,
and rotor speed [2,3]. CVD gives a means also for
reduction of hub losses by unloading the blade root
Corresponding author: Department of Fluid Mechanics (DFM),
Budapest University of Technology and Economics, Bertalan
Lajos u.4–6,Budapest H-1111, Hungary. email:
[4], and offers a potential to avoid highly twisted blades
[5]. Furthermore, in multi stage machinery, it provides
a strategy to realize an appropriate rotor exit flow angle
distribution [1].
Blade sweep, dihedral, and skew are known as tech-
niques of non-radial blade stacking. A blade has sweep
and/or dihedral if the sections of a datum blade of
radial stacking line are displaced parallel and/or nor-
mal to the chord, respectively [6]. A blade is swept
forward if the sections of a radially stacked datum
blade are shifted parallel to their chord in such a way
that a blade section under consideration is upstream
of the neighbouring blade section at lower radius
[3]. A special combination of dihedral and forward
sweep is referred to as circumferential forward skew
(FSK) [7,8]. In this case, the datum blade sections are
shifted in the circumferential direction, towards the
direction of rotation. By this means, the axial exten-
sion of the unskewed (USK) datum blading can be
retained, the blade mechanics is expected to be more
favourable than in the case of forward sweep alone,
JPE420 ©IMechE 2007 Proc. IMechE Vol. 221 Part A: J. Power and Energy
1012 J Vad, ARAKwedikha, Cs Horváth, M Balczó, M M Lohász, and T Réger t
and the following aerodynamic benefits, dedicated to
the incorporated forward sweep, can be utilized.
The open literature suggests a general consensus
that forward sweep/skew gives potential for the follow-
ing advantages in the part load operational range (flow
rates lower than design): improvement of efficiency,
increase of total pressure peak, and extension of stall-
free operating range by improving the stall margin
[3,712]. Nevertheless, the research results are rather
diversified regarding the judgment of performance
and loss modifying effects of forward sweep/skew at
flow rates near the design point. In reference [6], it
is pointed out generally that forward sweep near the
tip, i.e. ‘positive sweep’, gives a potential for reduction
of near-tip losses. Based on reference [9], application
of near-tip FSW can be recommended for efficiency
improvement over the operational range near the
design point [3,7,8], suggest that the application of
forward sweep along the entire span is beneficial for
loss reduction and performance improvement. How-
ever, forward sweep reported in references [10] and
[11] and FSK in reference [4] were found to cause the
deterioration of efficiency near the design point. In
reference [12], the reduction of efficiency was estab-
lished for a forward-swept rotor over the dominant
part of the entire stall-free operational range. Back-
ward sweep was reported to be optimal in reference
[13] from the viewpoint of efficiency improvement.
The performance and loss modifying aspects of
forward sweep/skew, which are specific to the indi-
vidual case studies as the above literature overview
suggests, are closely related to the three-dimensional
(3D) features of the blade passage flow [12,14,15].
Such 3D flow features are especially characteristic
for rotors of CVD, due to the spanwise blade circu-
lation gradient and the resultant vorticity shed from
the blade [2]. Although a number of reports are avail-
able on forward-swept and FSK rotors of CVD, e.g.
[4,79,16], no special emphasis is given to simul-
taneous application of CVD and non-radial blade
The current paper intends to present a case study
contributing to a more comprehensive understanding
of aerodynamic effects of FSK, CVD, and their combi-
nation, at the design flow rate. For this purpose, two
rotors of CVD, an USK and a FSK one, are aimed to
be compared qualitatively, by means of computational
fluid dynamics (CFD).
Rotor FSK under present investigation operates in
the open-type low-speed wind tunnel facility of the
Hungarian Institute of Agricultural Engineering (IAE),
Gödöll˝o, Hungary. The facility and the related custom-
built fan were designed at DFM, and were produced
by Ventilation Works Ltd., Hungary in 2004. The com-
ponents and instrumentation of the facility being
relevant to the present study are shown schematically
in Fig. 1. The main fan characteristics are summarized
in Table 1. Geometrical details of the rotor and out-
let guide vane (OGV) blading are specified in Table 2.
FSK was applied to the rotor blades in order to extend
the stall-free operating range. A virtual image of FSK,
obtained from the CFD technique, and a front-view
photo are presented in Fig. 2. The rotor and OGV
blade sections have C4 profiles [5,17] of 10 per cent
maximum thickness along the entire span, with circu-
lar arc camber lines. Results for a constant rotational
speed of 416 r/min are reported herein. The Reynolds
number, calculated with the blade tip circumferential
speed, the tip chord and the kinematical viscosity of
air at 20 C is approximately 1.074 ×106. The Mach
number which was computed with the blade tip cir-
cumferential velocity and the speed of sound in air
at 20 C is 0.13, and therefore, the flow is considered
Rotor FSK was originated from the virtual rotor
USK of radial stacking line, by shifting the blade sec-
tions of USK in circumferential direction towards the
direction of rotation, without making any modifica-
tions to the USK blade section geometry and stagger
angle distribution. The blade trailing edges (TEs) of
both USK and FSK fit to planes normal to the axis
of rotation. The skew angle in Table 2 is defined as
the angle between radial lines fitted to the TEs of
the datum and the shifted blade sections. The skew
angle is zero at the hub and increases along the span.
By this means, it was intended to avoid any stacking
line blend points, for which increased losses may be
expected [11]. Near the hub, the rotor blade sections
Fig. 1 Experimental facility and instrumentation (the
supporting struts for the nose cone and the hub
are omitted for simplicity)
Table 1 Main fan characteristics
Casing diameter 2000 mm D0.33
Hub-to-tip ratio ν0.600 D M FSK 0.27
Rotor blade count N12
OGV blade count 11
Tip clearance τ0.036
Proc. IMechE Vol. 221 Part A: J. Power and Energy JPE420 ©IMechE 2007
Aerodynamic effects of forward blade skew in axial flow rotors of CVD 1013
Table 2 Fan blading geometry
Rotor OGV
Fraction of span σ0 hub 0.25 0.50 mid 0.75 1.00 tip 0 hub 0.25 0.50 mid 0.75 1.00 tip
Solidity c/s1.38 1.01 0.89 0.80 0.72 1.93 1.50 1.32 1.19 1.18
Camber angle () 20.3 17.3 16.8 15.8 15.3 60.0 51.5 49.2 47.7 50.1
Stagger angle ()33.9 32.1 30.7 29.9 29.4 57.0 61.7 66.0 68.4 70.0
Skew angle () 0.0 0.0 0.3 1.6 3.5
Measured from circumferential direction.
Fig. 2 Virtual axonometric image and front-view photo
of FSK
are enlarged. This is favourable from the mechanical
point of view, and results in an aerodynamically ben-
eficial positive sweep and positive dihedral [6]atthe
blade root, as potential means of hub loss reduction.
In the following, ˆ denotes mass-averaging for ψid2 ,
ψ,ω, and ϕr, and area-averaging for ϕ. The USK rotor
is of CVD, i.e. the designed blade circulation increases
along the span, according to the following power
law [2,3]
ψid 2 D(R)=ˆ
ψid 2 D ) ·R
The CVD design concept was chosen in order to
make possible the preliminary design of each elemen-
tal blade cascade along the entire span using the same
cascade measurement data basis [17], and to reduce
blade twist and maintain chord length nearly constant
with span, for simplicity in manufacturing.
The experimental facility at IAE is not a test rig dedi-
cated for turbomachinery R&D; the FSK rotor under
investigation is its auxiliary unit. Consequently, the
facility is in absence of instrumentation expected in
turbofan studies. Nevertheless, it had been equipped
with an ad hoc, on-site measurement setup (Fig. 1),
in order to establish an experimental database for
validation of the CFD tool.
Characteristic curve and efficiency measurements
were carried out on the fan stage. The flow rate was
measured using the inlet bellmouth as an inlet cone,
calibrated on the basis of detailed velocity measure-
ments made in the test section. The total pressure
rise was considered as the difference of static pres-
sures measured downstream of the OGV and upstream
of the rotor in the annulus of constant cross-section
(equal upstream and downstream dynamic pressures
were assumed in the annulus). The differential pres-
sures playing role in the flow rate and total pressure
rise measurements were determined using Betz liquid
micromanometers. The constancy of rotor speed was
checked by means of a laser stroboscope.
The overall efficiency ηwas established as the ratio
between aerodynamic performance (product of vol-
ume flow rate and total pressure rise) and electric
power input to the frequency converter, measured by
a clamp meter. Although ηinevitably includes the
losses of the speed control unit, the electric motor, the
belt drive, and the bearings, it gives basic qualitative
information on the energetic behaviour of the fan.
Detailed flow velocity measurements were carried
out at the near-peak-efficiency point of =0.33, cor-
responding to the design flowrate. The velocity field
was measured using hot wire anemometry, in constant
temperature anemometer (CTA) mode, by means of a
DANTEC 9055P0511 type cross wire probe connected
to DISA 55M type CTA bridges equipped with servo
loop.The mobile CTA system is outlined in Fig. 1. vx1as
well as vx2and vu2 were measured along the radial span
having an axial position of 74.5 and 126.6 per cent
midspan axial chord, respectively, where the zero axial
position indicates the leading edge (LE) at midspan.
The radial traverses were carried out from 0.025 S
to 0.975 S, with resolution of 0.025 S. The sampling
rate provided 120 measurement readings per blade
passage at each radius along the circumference. The
measurements were taken at each radial position cov-
ering the progress of each blade passage 104 times. For
the CTA-based data presented herein, the velocity dis-
tributions representing the individual blade passages
have been circumferentially averaged.
Table 3 summarizes the pessimistically estimated
relative standard uncertainty of the measurement-
based quantities presented in the paper, at 95 per cent
confidence level, listing the most significant uncer-
tainty sources. The uncertainty analysis has been
carried out using the ‘root sum square’ method, follow-
ing the methodology in reference [12]. Any subvalue
JPE420 ©IMechE 2007 Proc. IMechE Vol. 221 Part A: J. Power and Energy
1014 J Vad, ARAKwedikha, Cs Horváth, M Balczó, M M Lohász, and T Réger t
Table 3 Experimental uncertainty
Quantity Source of uncertainty U(%)
/DUncertainty of inlet cone calibration ±1.5
Variation of operating state ±1.0
Uncertainty of differential pressure measurement ±0.5
/DOverall ±2.0
/DVariation of operating state ±1.0
Uncertainty of differential pressure measurement ±0.5
/DOverall ±1.2
DUncertainty of volume flowrate ±2.0
Uncertainty of total pressure rise ±1.2
Uncertainty of electric power measurement ±1.0
DOverall ±2.5
σUncertainty of measurement of endwall relative position ±0.5
ψid Uncertainty of adjusted volume flowrate ±2.0
Angular misalignment ±2.0
Temperature and pressure variation ±1.4
Uncertainty of velocity calibration ±1.4
Linearization error; voltage signal processing and A/D board resolution limits ±0.7
ψid Overall ±3.5
of Uin the table is not necessarily the error due to the
related uncertainty source in itself but the uncertainty
propagating due to this error (e.g. the Usubvalue
specified for the differential pressure measurement
for /Dis not the measurement uncertainty of the
manometer in itself). The overall uncertainties of the
quantities presented herein are taken as the square
root of sum of squares of Usubvalues. The uncertainty
is generally higher than expected in turbomachinery
studies [8], due to the ad hoc measuring technique
and to the non-laboratory environmental conditions.
The overall measurement uncertainty ranges are indi-
cated by error bars in the diagrams in the vicinity of
the measurement data points.
The flow fields in USK and FSK were simulated by
means of the commercially available finite-volume
CFD code FLUENT [18]. Referring to references [7], [8],
[16], and [19] reporting on computations for swept and
leaned fan and compressor rotors, the standard kε
turbulence model [20] has been used. The enhanced
wall treatment of FLUENT was applied, incorporat-
ing a blended model [21] between the two-layer
model and the logarithmic law of the wall. Among the
two-equation turbulence modelling options built into
FLUENT, this technique was found to give the most
reasonable agreement with the measurement results
presented later.
Taking the periodicity into consideration, the com-
putations regarded one blade pitch only. A typical
computational domain is presented in Fig. 3. The
domains extend to approximately 8 and 3.5 midspan
axial chord lengths upstream and downstream of the
rotor blading in the axial direction, respectively. The
Fig. 3 Computational domain for FSK (the casing is
hidden for clarity)
inlet face is a sector of the circular duct with 30central
angle. Downstream of the inlet face, sectors of the
steady inlet cone and the rotating hub with one blade
in the middle of the domain are included for both types
of blading.
At the inlet face, a swirl-free uniform axial inlet con-
dition corresponding to the actual flow rate has been
prescribed. The inlet turbulence intensity has been
set to 1 per cent, and the casing diameter was taken
as the hydraulic diameter for the calculation of the
turbulence length scale. Utilizing the features of the
annular cascade configuration, boundary conditions
of periodicity were applied. A zero diffusion flux con-
dition has been used for all flow variables at the outlet
boundary (outflow condition in FLUENT [18]).
Taking [19,22] as preliminary references, structured
hexahedral mesh has been developed for the entire
computational domain. This meshing technique is felt
promising from the viewpoint of computational accu-
racy. Furthermore, it offers a means to reduce the
computational cost by moderating the cell number.
About 50 per cent of the cells are located in the
refined domain in the vicinity of the blade. Taking
up the challenge of the relatively complicated blade
geometry, due to skew above midspan and LE sweep
Proc. IMechE Vol. 221 Part A: J. Power and Energy JPE420 ©IMechE 2007
Aerodynamic effects of forward blade skew in axial flow rotors of CVD 1015
Fig. 4 Finer mesh for FSK near the LE, TE, and tip
near the hub, the domain consists of 31 blocks. Fig. 4
shows representative segments and views of the mesh
for FSK near the LE, TE, and tip. An O-type mesh topol-
ogy has been built around the LE and TE, while H-type
topology is applied to the entire rotor blade passage.
Figure 5 presents a detail of the mesh topology in the
tip clearance region.
The equiangle skewness of a cell is defined as the
maximum value of the ratio of actual and possibly
highest deviation from the optimum angle, consider-
ing each vertex [18]. The grid design ensures that 99 per
cent of the cells have equiangle skewness less than 0.7,
and the maximum skewness value is 0.82. The highest
skewness values appear near the LE and TE. Over the
dominant part of the SS and PS, the skewness is less
than 0.25.
During the computations, the majority of y+values
fell within the range of 30–100, fulfilling the require-
ments of the applied wall law. The discretization of the
convective momentum and turbulent quantity fluxes
were carried out by the Quadratic Upstream Inter-
polation for Convective Kinematics (QUICK) method.
Fig. 5 Mesh topology in the tip clearance region
Typical computations required approximately 3000
iterations. The solutions were considered converged
when the scaled residuals [18] of all equations were
resolved to levels of order of magnitude of 106.
4.1 Grid sensitivity studies
Four discretization levels were used for the compu-
tation. Taking the ‘coarse’ mesh consisting of about
204.000 hexahedral cells as a basis, nearly uniform
refinement in axial, pitchwise, and spanwise direc-
tions resulted in the ‘mid’, ‘finer, and ‘finest’ meshes
(about 301.000, 494.000, and 694.000 cells, respec-
tively). The finer mesh, forming the basis of CFD
results presented in the paper, consists of 45 nodes
along the span. Clearance meshes resolved in span-
wise direction by 5, 9, and 17 nodes were tested, taking
the finer mesh as a basis. Application of nine nodes
in the clearance was concluded to be necessary, but
further refinement was found to be needless for the
fidelity of the numerical solution. For the finer mesh,
the outer domain (H-mesh) consists of 203, 27, and
54 grid nodes in axial, circumferential, and spanwise
directions, respectively.
The ideal total pressure rise was found to be the most
sensitive indicator of dependence of the numerical
solution on discretization. Figure 6 presents the ˆ
ψid 2
data computed for FSK at the design flow rate using
the four discretization levels. The grid-independency
of results based on the finer mesh is achieved on an
acceptable level from the aspect of present studies.
The computational data presented from this point
onwards are based on the finer mesh numerical
4.2 Validation analyses
Figure 7 shows the measured spanwise ˆϕ1,ˆ
ψid 2, and
ˆϕ2distributions established on the basis of CTA mea-
surement data for the design point. The experimental
data are compared in the figure with the distributions
Fig. 6 Influence of overall mesh refinement on the
numerical solution
JPE420 ©IMechE 2007 Proc. IMechE Vol. 221 Part A: J. Power and Energy
1016 J Vad, ARAKwedikha, Cs Horváth, M Balczó, M M Lohász, and T Réger t
Fig. 7 Measured and computed flow details for FSK.
Black dots: measurements, lines: CFD
computed for FSK at the axial positions of the
The ˆϕ1diagrams show the approximate realization
of the uniform axial rotor inlet condition used in
blade design. The computed ˆϕ1data fall below the
measured values near the hub, and the related ‘dis-
placement effect’ results in increased computed axial
velocity above midspan. The discrepancy of the near-
hub data is dedicated to the difference between the
realized and modelled inlet geometries, with special
regard to the inlet cone shape. Although the simula-
tion considers an inlet nose cone with smooth surface,
the inlet cone has eventually been assembled from
conical segments, as seen in Fig. 2, for manufacturing
simplification being accepted for industrial fans. The
edges appearing at the connection of the segments
act as turbulence generators, refreshing the hub inlet
boundary layer otherwise being thickened.
The rotor inlet axial velocity underpredicted by CFD
leads to higher flow incidence and blade load (lift)
below midspan. Considering nearly unchanged free-
stream relative velocity w, the increased blade lift of
an elemental cascade leads to increased outlet swirl
and ideal total pressure rise, according to the following
classic approximate relationship [5,11,17], assuming
swirl-free inlet far upstream
Just the opposite tendency, i.e. decreased incidence,
lift, outlet swirl, and ideal total pressure rise is expected
above midspan where the rotor inlet axial velocity
is overpredicted in comparison with the measure-
ments. The trends explained above appear in the
ψid 2 plots where the computed data are higher and
lower than the measurement-based ones below and
above midspan, respectively. The ‘theoretical’ ˆ
ψid 2
distributions specified in Fig. 7, calculated from 20
to 80 per cent span using the model described in
Appendix 2, correlate fairly well with the CFD as
well as with the measurement-based ˆ
ψid 2 diagrams.
This confirms the physical relevance and consistency
of both the measured and computed ˆϕ1,ˆ
ψid 2, and ˆϕ2
data sets.
Besides the above described incidence effect,
another reason for the discrepancies above midspan,
especially near the tip, is the limited capability of the
applied turbulence model. However, even with the
presence of the incidence effect, the relative differ-
ences between the computed and measured ˆ
ψid 2 and
ˆϕ2data reported here do not exceed, up to 90 per
cent span, the maximum differences valid for a rep-
resentative forward-skewed fan (AV30N fan, 30FSK)
studied in references [7] and [8] involving standard
kεmodelling. It should be noted that the validity of
the CFD technique in references [7] and [8] has been
accepted for widespread investigation of CVD rotors
with non-radial blade stacking.
All of the qualitative features judged to be essential
for the validity of the CFD tool on the basis of refer-
ence [3] – i.e. the overturning (increased ˆ
ψid 2)near the
rotating hub; the spanwise increase of ideal total pres-
sure rise, fitting to the CVD concept [2]; the peak in
ψid 2 near the blade tip due to the presence of high-
loss fluid; and the decrease of swirl near the casing
due to the underturning effect of the stationary cas-
ing wall and the leakage flow – are resolved by the
The validity of the CFD method enables the rep-
resentation of the following trends observed in the
measured ˆϕ2data: axial velocity reduction near the
blade root due to the hub boundary layer; increasing
Proc. IMechE Vol. 221 Part A: J. Power and Energy JPE420 ©IMechE 2007
Aerodynamic effects of forward blade skew in axial flow rotors of CVD 1017
axial velocity along the dominant part of span, due to
the CVD concept [2,16]; and velocity defect near the
casing, due to the presence of high-loss fluid as well
as the casing boundary layer and leakage flow. The ˆϕ2
values below midspan and the predicted location and
value of maximum axial velocity are in fair agreement
with the experiments.
The measured and computed characteristic and effi-
ciency curves are shown in Fig. 8. CFD has been
calculated on the basis of the difference between
the computed mass-averaged static pressures at the
rotor outlet and inlet CTA measurement locations in
the annulus. The total pressure and flow coefficient
data are normalized by the corresponding values of
the measured FSK design point (D M FSK =0.27, D=
0.33). ηCFD was calculated as the product of com-
puted global total pressure rise and volume flow rate
data divided by the computed shaft power input. The
efficiency data have been normalized by appropriate
reference values taken at the design flow rate. Polyno-
mial trend lines have been fitted to the data points in
the figure.
The [D,D M FSK] design point and the slope of
the M FSK() curve near the design flow rate are
fairly well captured by the simulation. The measured
and computed trends of efficiency variance from the
Fig. 8 Measured and computed global performance
design point towards moderately lower flow rates are
also in fair agreement.
5.1 Comparison of USK and FSK performance
Figure 8 offers a comparison between the performance
curves computed for USK and FSK. Despite the limited
capability of the applied turbulence model at lower
flow rates, the computed () curves represent the
following well-known qualitative features dedicated
to forward sweep/skew: (a) if no blade correction is
applied for retaining the original Euler work, is
reduced near the design flow rate [4,7,8,12,14,16],
(b) the total pressure peak is shifted towards lower flow
rates, and (c) is improved at flow rates considerably
lower than the stall margin of the rotor with radially
stacked blades [3,11]. The computed η() plots show
that the deterioration of total efficiency is less drastic
for FSK when throttling from the design flowrate.
The total efficiency computed for FSK at the design
point falls below the value for USK. This observa-
tion, fitting to former experiences in references [4] and
[1012], is the aspect provoking the discussion in the
following sections.
5.2 Design flowrate: pitchwise averaged data
Figure 9 presents the spanwise distribution of pitch-
wise averaged values for the dimensionless rotor inlet
and outlet axial velocities as well as radial velocity,
ideal total pressure rise, and total pressure loss coeffi-
cient at the outlet. The inlet (‘1’) and outlet (‘2’) planes
have the axial position of 20.0 and 113.0 per cent
midspan axial chord, respectively, where the zero axial
position indicates the LE at midspan.
As the figure suggests, the applied blade skew has
an influence on the rotor inlet flow field: the inlet axial
velocity for FSK is increased near the tip and is reduced
at lower radii, as can be observed for FSK rotors in ref-
erence [4]. The outlet axial velocity is increased below
midspan for FSK. The difference in radial rearrange-
ment of fluid for USK and FSK, i.e. radially inward
dominant flow for FSK [4,7], is visible on the outlet
radial velocity plots. As the ideal total pressure rise
and axial velocity plots show, FSK performs increased
and decreased Euler work compared to USK below and
above midspan, respectively. Such trend appears in
reference [7] as well (AV30N fan). The Euler work at
the tip is reduced due to non-radial blade stacking, as
was observed in [11].
Figure 9 presents also the ˆ
ψid 2 D and ˆϕ2D distribu-
tions that were determined as outlined in Appendix 2.
These distributions indicate the increase of ˆ
ψid 2 D and
JPE420 ©IMechE 2007 Proc. IMechE Vol. 221 Part A: J. Power and Energy
1018 J Vad, ARAKwedikha, Cs Horváth, M Balczó, M M Lohász, and T Réger t
Fig. 9 Pitchwise averaged data. White dots: USK, black
dots: FSK
ˆϕ2D along the span due to the CVD concept. They
served as a basis for the preliminary design of the rotor
blade sections further from the annulus walls. Con-
sidering the non-uniformity of CFD-predicted axial
rotor inlet condition, which differs from that used
in the design concept, the agreement between the
design and USK distributions is fair farther from
the endwalls. However, increased discrepancy can be
observed between the design and FSK distributions.
Although the total pressure loss is reduced near the
tip, it is increased over the dominant part of span due
to skew. The same tendency was reported in reference
[11] for a rotor with forward sweep at the tip.
The above tendencies will be explained in the fol-
lowing section, by means of analysis of the detailed
flow field. Rotor inlet and outlet flow maps will be pre-
sented. Furthermore, the flow field will be surveyed at
20 and 90 per cent span, being two representative loca-
tions where significant differences occur in the fluid
mechanical behaviour of USK and FSK (Fig. 9).
5.3 Design flowrate: pitchwise resolved data
Figure 10 presents the maps of ideal total pressure rise,
axial and radial velocities, and total pressure loss coef-
ficient at the rotor outlet. The regions downstream of
the SS and PS, separated by the blade wake zone, are
indicated by appropriate labels. These data reflects
the trends seen in Fig. 9. For USK, spanwise increase
of ψid 2 dominates along the span, according to the
CVD concept based on equation (1). The spanwise
gradient of Euler work and blade circulation results
in increasing axial velocity along the dominant part of
the span according to the physical concept described
in Appendix 2, and in vortices shed from the TE. The
TE shed vorticity induces radially inward and outward
flow on the PS and SS, respectively, as observed also in
references [2] and [3].
Circumferential FSK causes substantial changes in
the 3D blade passage flow structure.The spanwise gra-
dient of ψid 2 is reduced for FSK, for reasons explained
later. This trend was observed also in references [4]
and [7]. Based on the physical principle expressed in
equation (7) in Appendix 2, the moderation of span-
wise ψid 2 gradient causes the moderation of spanwise
variance of ϕ2. The theoretical ˆϕ2plots in Fig. 9, com-
posed as described in Appendix 2, and correlating
fairly well with the CFD data, justify this physical trend.
The reduction of d ˆϕ2/dRcorresponds to an increase
and a decrease of ϕ2below and above midspan,
respectively, as was found also in references [4] and
[79]. According to continuity, this yields the domi-
nance of inward flow in terms of pitchwise averaged
radial velocity (negative ˆϕr2 values for FSK in Fig. 9),
corresponding to the amplification and the attenua-
tion of radially inward and outward flow on the PS
and SS, respectively. The moderation of d ˆ
ψid 2/dR,
Proc. IMechE Vol. 221 Part A: J. Power and Energy JPE420 ©IMechE 2007
Aerodynamic effects of forward blade skew in axial flow rotors of CVD 1019
Fig. 10 Outlet flow maps. Left column: USK, right
column: FSK
i.e. reduction of spanwise blade circulation gradient,
results in the attenuation of TE shed vorticity [2,5]
for FSK, also contributing to the moderation of radial
outward flow on the SS.
The mechanism by which FSK attenuates the SS
radial outward flow is demonstrated in Fig. 11. Due to
FSK, the isobars in the decelerating region are inclined
‘more forward’ for FSK than for USK. Therefore, the
local radial outward flow is moderated, the flow is
guided ‘more inward’ for FSK on the SS. Such radial
flow controlling effect has been described qualitatively
in reference [9].
The moderation of d ˆ
ψid 2/dRdetected for FSK is
explained as follows. Figure 12 shows the axial velocity
and ideal total pressure rise maps at the rotor inlet.
The upstream regions where the forward effect of SS
Fig. 11 Distribution of static pressure coefficient Cpon
the SS Left: USK, right: FSK
and PS phenomena can be detected are indicated
by appropriate labels. A zone of pronounced suc-
tion effect can be observed upstream of the SS of
the near-tip region of FSK, indicated by increased
axial velocity and counter-swirl compared with USK.
Upstream of the PS of FSK, locally reduced axial veloc-
ity and increased swirl appear, compared with USK.
Pitchwise-averaging points out that ˆϕ1(Fig. 9) and
the Euler work are higher for FSK near the tip at the
rotor inlet. This is suggested also by the generally
increased ψid and ϕdata near the FSK LE in Fig. 13.
The reason for the above-mentioned is that the near-
tip part of the forward-skewed blade protrudes into
the upstream relative flow field, and carries out work
in advance compared to the blade sections at lower
radii. According to the conservation of mass at the pre-
scribed design flowrate, increase of inlet axial velocity
near the tip results in the reduction of inlet axial veloc-
ity at lower radii of FSK, as was already indicated in
Fig. 9. The reduced axial velocity results in increased
flow incidence angle, manifesting itself in increased
lift, i.e. increased depression and overpressure on the
SS and PS, respectively. This is illustrated in the Cp
Fig. 12 Inlet flow maps. Left column: USK, right
column: FSK
JPE420 ©IMechE 2007 Proc. IMechE Vol. 221 Part A: J. Power and Energy
1020 J Vad, ARAKwedikha, Cs Horváth, M Balczó, M M Lohász, and T Réger t
Fig. 13 Flow characteristics at 90 per cent span
plots of Fig. 14. As equation (2) suggests, the higher lift
being valid for FSK at lower radii potentially leads to
increased Euler work and blade section performance.
Indeed, as Fig. 14 indicates, FSK performs higher ideal
total pressure rise and axial velocity at lower radii,
compared with USK, as was suggested already in Fig. 9.
Figure 13 shows increased loss on the SS of FSK
near the tip, for the following presumed reason. Cir-
cumferential FSK results in positive sweep [6] near
the tip, with leakage loss-reducing effects anticipated,
but inevitably also in negative dihedral, i.e. acute
angle between the suction surface and the casing
wall. As presumed on the basis of reference [6], nega-
tive dihedral results in increased near-tip and leakage
losses. The unfavourable effect of negative dihedral
appears to dominate over the favourable effect of pos-
itive sweep from the viewpoint of losses near the tip,
although the tip sweep angle is considerably larger
than the tip dihedral angle (approximately 22and
13, respectively).
As the ω2plots in Fig. 10 suggest, blade sections of
FSK away from the tip also have increased loss on the
SS. This is mainly due to the increased flow incidence
Fig. 14 Flow characteristics at 20 per cent span
angle and the resultant higher adverse pressure gradi-
ent. The increase of losses further from the endwalls
in a rotor of forward-swept tip was also discussed in
reference [11] to the unfavourable conditions in the
SS boundary layer.
Comparative CFD studies have been carried out on
two rotors – USK and FSK – at the design flow rate,
in order to investigate the aerodynamic effects of
CVD and circumferential FSK, without geometrical
correction of the elemental blade cascades of the
skewed blading. Preliminary studies were published in
reference [23].The results are summarized as follows.
1. The studies indicated that the circumferentially
forward-skewed blade tip carries out work on the
incoming fluid in advance compared with the blade
sections at lower radii, due to its protrusion into
the upstream relative flow field. This results in
increased and decreased inlet axial velocities near
the tip and at lower radii, respectively.
Proc. IMechE Vol. 221 Part A: J. Power and Energy JPE420 ©IMechE 2007
Aerodynamic effects of forward blade skew in axial flow rotors of CVD 1021
2. The decreased axial velocity at lower radii leads to
increased incidence, lift, and blade performance.
Such uploading below midspan, coupled with
unloading above midspan due to sweep, reduces
the spanwise gradient of Euler work. Consequently,
the blade circulation and axial velocity distribution
prescribed along the span by the CVD concept
tends towards that of a free vortex flow pattern. This
results in the decrease of global ideal total pressure
3. Increased total pressure loss was found along the
dominant portion of the span of FSK. This was
dedicated to (a) the negative dihedral near the
tip, always incorporated by circumferential FSK,
and (b) predominantly due to the off-design cas-
cade conditions at lower radii, i.e. increased flow
incidence due to the tip forward effect, and the
related higher SS adverse pressure gradients. Due
to the reduced global ideal total pressure rise and
increased losses, the global total pressure rise and
total efficiency of FSK were found to be reduced
compared with the USK rotor.
4. For rotors of CVD, the radial outward flow on the
SS is intensified in comparison with free vortex
rotors, due to the vortices shed from the TE in
accordance with the spanwise increasing blade cir-
culation. This suggests that in CVD bladings, the SS
boundary layer fluid has increased inclination to
migrate outward and to accumulate near the tip. As
the present studies indicated, forward sweep atten-
uates the radial outward flow on the SS. This yields
that the application of forward sweep for potential
reduction of near-tip loss is especially welcome for
CVD rotors.
5. The present study, supplemented with literature
data cited in the introduction, suggests that appli-
cability of ad hoc blade stacking techniques is
doubtful in the achievement of efficiency gain
and prescribed performance at the design flow
rate. Instead, application of reliable CFD-based
design systems [13] is recommended for systematic
consideration and control of both load- and loss-
modifying effects due to non-radial blade stacking.
This work has been supported by the Hungarian
National Fund for Science and Research under con-
tracts No. OTKA T 043493 and K63704, and, on the
behalf of Cs. Horváth, out of the József Öveges Pro-
gram HEF_06_3 (BMEGPK06). Gratitude is expressed
to Prof László Fenyvesi and Mr József Deákvári, Hun-
garian IAE, Gödöll˝o, for contributing to the measure-
ments, and to Mr Lóránt Farkas, Szell˝oz˝oM˝uvek Kft.
(Ventilation Works Ltd), for consultation.
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cblade chord
CLblade lift coefficient
Cplocal static pressure coefficient
ref /2)
Mexponent in the design power law,
equation (1)
nrotor speed
Nrotor blade count
pstatic pressure
rradius =d/2
Rdimensionless radius =r/rt
sblade spacing (blade pitch) =dπ/N
Sblade span =(dtdh)/2
trotor tip clearance
uref reference velocity =dtπn
Urelative standard experimental
vflow velocity in the absolute frame of
wrelative free-stream velocity
y+wall normal cell size (in wall units)
ptlocal total pressure rise
ηglobal total efficiency
ηoverall efficiency
νhub-to-tip ratio =dh/dt
ρfluid density
σfraction of span (radial distance from
the hub divided by S)
τrelative tip clearance =t/S
ϕlocal axial flow coefficient =vx/uref
ϕrlocal radial flow coefficient =vruref
global flow coefficient (annulus area-
averaged axial velocity divided by uref )
global total pressure coefficient
(annulus mass-averaged total
pressure rise divided by ρuref 2/2)
ψlocal total pressure rise coefficient
=pt/uref 2/2)
ψid local ideal total pressure rise coef-
ficient =ptid/ u2
ref /2)=2Rvu/uref
(from the Euler equation of turboma-
chines, considering swirl-free inlet far
ωtotal pressure loss coefficient
=ψid ψ
Subscripts and superscripts
CFD based on CFD data
D design; at the design flow rate
FSK circumferentially forward-skewed
h hub
id ideal (inviscid)
M based on measurement data
r radial coordinate
t blade tip
u tangential coordinate
USK unskewed blading
xaxial coordinate
1 rotor inlet plane
2 rotor exit plane
ˆpitchwise averaged value
passage-averaged value
Calculation of approximate theoretical spanwise
distributions of flow characteristics
Pitchwise averaged quantities are considered herein.
The superscript ˆhas been omitted for simplicity.
At a given radius, the total pressure rise realized by
the rotor is
pt=ηptid =pt2 pt1 =p2+ρv2
Proc. IMechE Vol. 221 Part A: J. Power and Energy JPE420 ©IMechE 2007
Aerodynamic effects of forward blade skew in axial flow rotors of CVD 1023
The following simplifying assumptions are taken.
1. The flow is incompressible, i.e. ρ=constant.
2. Although the local total efficiency in equation (3)
varies along the span, it is assumed to be constant
farther from the endwalls at the design flowrate, on
the basis of measurement data in [12].
3. The inlet swirl is neglected, i.e. vul=0, and the
streamlines are parallel upstream of the rotor,
i.e. the normal component of Euler equation
in the natural coordinate system reads p1(r)=
4. The radial velocity components are neglected, i.e.
x1and v2
Taking the radial derivative of equation (3), and
applying the above simplifications, reads
The Euler equation of turbomachines for swirl-free
inlet is as follows
ptid =ρuvu2 (5)
According to the Euler equation, dp2/dris expressed as
Substituting equations (5) and (6) to equation (4),
rearranging, and putting into a dimensionless form
dψid 2
dRηψid 2
When determining the theoretical ψid 2(R)distribu-
tions in Fig. 7, the measured as well as the computed
ϕ1(R)and ϕ2(R)distributions were approximated as
linear functions from 20 to 80 per cent span, using the
least squares method. This provided for local approx-
imate data of ϕ1,dϕ1/dR,ϕ2, and dϕ2/dRto be substi-
tuted into equation (7). The differential equation (7)
was solved for ψid 2(R)numerically for the spanwise
region of axial velocity linearization, retaining the
computed ψid 2 data at midspan as boundary condi-
tion. For determination of the theoretical ϕ2(R)distri-
butions in Fig. 9, the ϕ1(R)and ψid 2(R)distributions
were linearized, and equation (7) was solved numer-
ically, retaining the computed ϕ2data at midspan as
boundary condition. η=0.90 was set for each case as
representative value, based on reference [12].
The ψid2 D(R)and ϕ2D(R)distributions shown in Fig. 9
were determined on the basis of equations (1) and
(7), but assuming uniform axial inlet condition, apply-
ing empirical corrections considering the spanwise
change of efficiency and the blockage due to the annu-
lus wall boundary layers, and taking the prescribed D
and Ddata as integral conditions.
JPE420 ©IMechE 2007 Proc. IMechE Vol. 221 Part A: J. Power and Energy
... where and a are the kinematic viscosity and the speed of sound of the gas, respectively. The hub-to-tip diameter ratio j is defined by Fig. 1 Examples of low-pressure axial fans: (a) large hub-to-tip ratio fan [3] and (b) small hub-to-tip ratio fan [4] where D hub is the hub diameter. In Fig. 1, examples are shown of fans with a large and a small hub-to-tip ratio j. ...
... For fans with a fairly large hub-to-tip ratio j ¼ 0:6 (and including diffuser blades), the influence of forward skewed blades with a prescribed vortex distribution on the aerodynamic performance of axial fans has been investigated in Refs. [3] and [22] by experiments and CFD simulations (using the k À turbulence model). The CFD simulations account for the tip gap and the mesh independence of the CFD results has been checked. ...
... Besides measured velocities (using hot-wire anemometry) at the design point, the aerodynamic performance has also been estimated. The agreement between measured and computed performance characteristics is labeled fair (10% deviation at the design point) due to limitations in the experimental test setup [3]. ...
Full-text available
Rotor-only ducted axial fans with small hub-to-tip diameter ratio are widely used in many branches of industry, especially for cooling and ventilation purposes. For such fans, extensive regions of backflow are present downstream of the fan near the hub. Only few Computational Fluid Dynamics (CFD) studies for such fans have been reported in the scientific literature. In order to develop guidelines for obtaining accurate CFD predictions for such fans, validation simulations of a fan with small hub-to-tip diameter ratio have been performed by comparing experimental and computed aerodynamic performance characteristics. These guidelines pay special attention to the trailing edge shape, presence of non-aerodynamically shaped blade sections, tip gap and employed turbulence model. The results \MOD{for the fan studied here} show that the actual (rounded) trailing edge is necessary; the main blade (without non-aerodynamically shaped blade sections) well represents the aerodynamic performance of the whole fan blade; it is recommended not to take the tip gap into consideration due to the existence of significant flow separation; the use of the Spalart-Allmaras turbulence model is advised for giving better agreement with measurements.
... Good agreements were found in velocity distribution at the fan outlet with measurements and flow fields were further analyzed in CFD results. J. Vad [2] in 2007 and Jin [3] in 2009 investigated aerodynamic effects of blade skew in axial fan by CFD simulations, the total pressure rise coefficient were used to validate the prediction in both investigations and fairly good agreement was found. In 2018, the effects of Gurney flap, winglets assembled on fan blade were investigated respectively by Liu Chen [4] and Jae Hyuk Jung [5] by CFD simulations. ...
... These investigations prove that the CFD indeed became a powerful tool for axial fan research, however, the simulation strategies of these investigations are different. At the same time, although CFD has been applied successfully in many axial fan investigations, the simulations with low (a) High Hub-to-tip Ratio Fan [2] (b) Low Hub-to-tip Ratio Fan [7] ...
... For the CFD simulations of low-pressure axial fan, some researches [1,15] ignore the tip gap, but in some investigations the tip gap is considered [2,4], so it is not clear whether the tip gap should be taken into account as simulation strategy for axial fan with low hub to tip ratio. The simulations with and without tip gap are performed to investigate the effects of tip gap on the simulation predictions. ...
... The estimation of the absolute errors followed the methodology in Ref. [19]. The errors for the hot-wire-measured velocity data were estimated on the basis of Ref. [20], considered as a departmental preliminary study to the application of hot-wire anemometry to flow phenomena related to axial flow fans. ...
... (19)) can be explored, theoretically obtained by simultaneous fulfillment of conditions in Eqs. (18), (20), (21), (26), and (27); or in Eqs. (18), (21), (29), and (30). ...
The paper presents hot wire measurements in a wind tunnel, close downstream of basic models of blade sections being representative for low-speed, low-Reynolds-number axial fans, in order to explore the signatures of vortex shedding (VS) from the blade profiles. Using the Rankine-type vortex approach, an analytical model was developed on the velocity fluctuation represented by the vortex streets, as an aid in evaluating the experimental data. The signatures of profile VS were distinguished from blunt-trailing-edge VS based on Strouhal numbers obtained from the measurements in a case specific manner. Utilizing the experimental results, the semiempirical model available in the literature for predicting the frequency of profile VS was extended to low-speed axial fan applications. On this basis, quantitative guidelines were developed for consideration of profile VS in preliminary design of axial fans in moderation of VS-induced blade vibration and noise emission.
... The most striking features of this low-speed axial fan are its swept and skewed blade shape [13]. Previous studies have demonstrated that the forward-tilted blade reduces the total pressure loss near the rotor hub and expands the stable working range of the fan [14][15][16]. Meanwhile, sweeping the blades backward will increase the air velocity radial component, incrementing the downstream turbulence energy and fan noise [17,18]. ...
Full-text available
Outer edge bending is already used on the axial fan blades of air conditioners, reducing the leakage flow loss at the blade tip and suppressing the tip vortex development, thereby improving fan aerodynamic and acoustic performance. However, there are few studies on the multi-parameter design and optimization of this complicated structure, and most studies only focus on the overall sound pressure level rather than the sound quality when evaluating the fan noise. This study investigated the effects of outer edge bending structure on the aerodynamic performance and sound quality of air conditioners’ axial fans by experiments and numerical methods. Based on the orthogonal design method, the effects of three bending parameters, the circumferential starting angle, radial relative position, and the bending degree effects on the performance of the axial flow fan blade were analyzed, and the best efficiency scheme was selected. A comparative analysis of the preferred and the original bending schemes shows that the bending towards the blade suction surface successfully inhibits the development of tip leakage vortex at the blade tip, thereby achieving efficiency enhancement and noise reduction. The experimental results show that the preferred bending scheme with a 10° circumferential starting angle, 90% radial relative position, and 8% bending degree can effectively reduce the fan’s broadband noise within 200~1000 Hz by 0.54~2.68 dB (A) at different operating conditions. Additionally, the preferred bending blade with reasonably designed bending effectively reduced the loudness and roughness of the fan noise in the rated conditions, and the sound quality of the studied fan was correspondingly improved.
... Better performance in the reduction of noise [17e20] and total pressure loss [20] have been documented in published literature. Efficiency improvement is also highlighted in papers like [19,21,22]. In the flow velocity distribution optimization, the skewed blade principle works on defining the effective blade exposure to the air. ...
In axial ventilation fans, the generation of a uniform flow velocity is desirable for better efficiency. To that end, different fan blade types have been developed to achieve better flow uniformity. This article aimed to characterize the flow distribution and its uniformity in four blade designs, namely constant chord, tapered blade, skewed blade, and tapered skewed blade, using Computational Fluid Dynamics (CFD). The study employs an iterative study where key study decisions are made as the study progresses. The study began with the selection of a blade profile for the study. A comparative study between the NACA seven-digit and four-digit series was conducted and for its higher flow throughput, the four-digit airfoil profile was selected. Next, with 30 and 40° Angle of Attack (AoA), the constant chord blade flow pattern is characterized. At 40° AoA flow disturbance and high-velocity spots were observed establishing the problem statement. Following that, three optimization strategies (tapering, skewing, and taper skewing) were applied in the design, and the flow pattern of each design was studied. Using a dispersion study a flow uniformity comparison between the models conducted. The property trade-off between three key performance indicators: efficiency, flow rate, and flow uniformity studied. The result shows an axial fan having a higher efficiency doesn't necessarily mean it has higher throughput whereas lower flow dispersion relates to the system's higher efficiency. Therefore, it can be concluded that seeking higher efficiency and flow uniformity in the design and development of axial fans comes with system throughput trade-off.
... The peak load of the fan blades is reduced, as the fan-blade leading edge is no longer centric, which leads to weaker interaction forces of the fan blade with the inflow [9,10]. The type of skew also changes the loading along the fan span: forward skew is known to decrease the blade loading in the tip region and backward skew to increase the loading [11][12][13][14][15]. Additionally, fan-blade skew leads to a reduction of turbulence ingestion noise compared with fans without any fan-blade skew [4,[16][17][18]. ...
Axial fans represent a major source of noise in technical systems. Based on investigations on airfoils, a promising measure for reducing the sound emission of such systems with axial fans is the application of leading-edge serrations to the fan blade. Hence in this study, the joint impact of fan-blade skew—as a commonly used noise-reduction approach—and leading-edge serrations on the sound emission of low-speed axial fans was investigated. Forward-, backward-, and unskewed fans with five different leading-edge modifications were used. For each fan, a reference configuration with straight leading edges and four configurations with sinusoidal leading-edge serrations (with two different amplitude values and two different wavelength values) were examined. The results show that, among the reference fans with unmodified leading edges, the forward-skewed fan had the lowest sound emission. The serrations applied to the unskewed fan lead to an increase in the efficiency and a decrease in the sound emission. When applied in combination with fan-blade skew, the serrated fans showed an even lower sound emission and a higher efficiency—even the forward-skewed fan, which already has a low sound emission without any leading-edge modifications. The highest sound reduction was achieved by applying the serrations with the smallest wavelength and the highest amplitude. The findings prove that even for low-noise fans with forward-skewed fan blades, a further sound reduction is possible with the use of leading-edge serrations.
... Furthermore, the increase in the stall margin of the impeller due to the forward-skewed blades is greater than that of the backward-skewed blades. Vad et al. [5] designed circumferentially skewed blades of an axial fan, and their experimental results showed that the forward-skewed blades enhance the performance of blades by increasing the axial velocity near the tip and by reducing the axial velocity near the lower half of the blade. However, both the total pressure rise and the efficiency decrease. ...
Full-text available
For a single-stage variable-pitch axial fan, the aerodynamic performance and through flow with and without blade skewing are examined numerically. Simulated results show that the total pressure rise and efficiency increase by 2.99% and 0.16%, respectively, with the best forward-skewed angle of θ = 3° at the design conditions. At the blade pitch angles of β = 29° and 35°, the total pressure rises and efficiency of the fan with θ = 3.0° under the highest efficiency point change by −0.55%, −0.53% and 1.39%, 2.11%, respectively. At design and off-design conditions, the forward-skewed blades mitigate tip leakage and delay the emergence of separation flow at the blade root, these benefits are higher at the higher blade pitch angle. The θ = 3.0° forward skew effectively raises the stage performance of the impeller and guide vanes.
... The peak load of the fan blades is reduced, as the fan-blade leading edge is no longer centric, which leads to weaker interaction forces of the fan blade with the inflow [9,10]. The type of skew also changes the loading along the fan span: forward skew is known to decrease the blade loading in the tip region and backward skew to increase the loading [11][12][13][14][15]. Additionally, fan-blade skew leads to a reduction of turbulence ingestion noise compared with fans without any fan-blade skew [4,[16][17][18]. ...
Compressor one-dimensional design prediction models provide match basis of cascade design parameters. Solidity and aspect ratio are two primary prediction parameters. Compound lean blade is a three-dimensional blade technology generally applied in compressor redesign. This paper analyzes the correlations of solidity, aspect ratio and compound lean blade in order to find compound lean blade effect on cascade design. Abundant cascade cases have been computed by proved numerical method. Characteristics of compound lean blade effect on diffusion and loss are discussed. To exploit the dependence of geometry and performance parameters under min-loss working condition, a minimum loss incidence (io) model, a diffusion factor (Do) model and a total pressure loss (ωo) model are built up, and dependence map of parameters is analyzed. It suggests that compound lean blade changes constraint relations between parameters and enriches the choices of design case. Finally, a minimum loss model (ωθ,Do)min and a minimum blade count model (h/tθ,Do)min are studied, providing new parameter match principles and selection range, which breaks through traditional parameters balance relationship established based on straight blade. So, there is reasonable prediction that introducing compound lean blade to prediction models could broaden compressor one-dimensional design view and contribute to compressor optimal design.
Conference Paper
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Shape optimization of a transonic axial compressor rotor operating at the design flow condition has been performed using the response surface method and three-dimensional Navier-Stokes analysis. The three design variables, blade sweep, lean and skew, are introduced to optimize the three-dimensional stacking line of the rotor blade. The objective function of the shape optimization is adiabatic efficiency. Throughout the shape optimization of the rotor, the adiabatic efficiency is increased by reducing the hub corner and tip losses. Separation line due to the interference between a passage shock and surface boundary layer on the blade suction surface is moved downstream for the optimized blade compared to the reference one. Among the three design variables, the blade skew is most effective to increase the adiabatic efficiency in the compressor rotor.
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Experimental studies were carried out in order to survey the performance and efficiency aspects of spanwise constant forward and backward sweep in axial flow rotors of low aspect ratio (AR) blading for incompressible flow, for part-load, near-design, and overload operational ranges. The experiments involved overall performance tests, laser Doppler velocimetry, and stationary total pressure probe measurements. The overall performance data and pitchwise averaged and resolved flow characteristics were evaluated in detail. For moderate and high flow rates, it was pointed out that positive or negative sweep tends to reduce or increase the blade load in the vicinity of the endwalls, respectively. It has been concluded that the loss-modifying effect of sweep can be judged by considering the three-dimensional viscous phenomena, and the influence of sweep on local blade efficiency depends on the balance of changes in blade load and losses. For low flow rates, forward sweep was found beneficial over the entire span from the aspect of improved stall margin and efficiency. The influence of AR on the performance reducing effect of sweep was studied on the basis of literature data.
Conference Paper
The paper presents an analytical model for qualitative prediction of physical effects stimulating or retarding the radial outward migration of high-loss fluid in the suction side boundary layer of the blades of isolated axial flow rotors. This fluid motion is considered as the main cause of tip stalling. In elaboration of the model, the radial acceleration of fluid particles is expressed, and the Reynolds-averaged Navier-Stokes equation is applied for description of flow. On the basis of the model, guidelines are formulated in order to retard the inception of stall.
Conference Paper
Experimental and computational studies were carried out in order to survey the energetic aspects of forward and backward sweep in axial flow rotors of low aspect ratio blading for incompressible flow. It has been pointed out that negative sweep tends to increase the lift, the flow rate and the ideal total pressure rise in the vicinity of the endwalls. Just the opposite tendency was experienced for positive sweep. The local losses were found to develop according to combined effects of sweep near the endwalls, endwall and tip clearance losses, and profile drag influenced by re-arrangement of the axial velocity profile. The forward-swept bladed rotor showed reduced total efficiency compared to the unswept and swept-back bladed rotors. This behavior has been explained on the basis of analysis of flow details. It has been found that the swept bladings of low aspect ratio tend to retain the performance of the unswept datum rotor even in absence of sweep correction.
The current paper reports on investigations with an aim to advance the understanding of the flowfield near the casing of a small-scale high-speed axial flow compressor rotor. Steady three-dimensional viscous flow calculations are applied to obtain flowfields at various operating conditions. To demonstrate the validity of the computation, the numerical results are first compared with available measured data. Then, the numerically obtained flowfields are analysed to identify the behaviour of tip-leakage flow, and the mechanism of blockage generation arising from flow interactions between the tip clearance flow, the blade/casing wall boundary layers, and non-uniform main flow. The current investigation indicates that the ‘breakdown’ of the tip-leakage vortex occurs inside the rotor passage at the near stall condition. The vortex ‘breakdown’ results in the low-energy fluid accumulating on the casing wall spreads out remarkably, which causes a sudden growth of the casing wall boundary layer having a large blockage effect. A low-velocity region develops along the tip clearance vortex at the near stall condition due to the vortex ‘breakdown’. As the mass flow rate is further decreased, this area builds up rapidly and moves upstream. This area prevents incoming flow from passing through the pressure side of the passage and forces the tip-leakage flow to spill into the adjacent blade passage from the pressure side at the leading edge. It is found that the tip-leakage flow exerts a little influence on the development of the blade suction surface boundary layer even at the near stall condition.
The paper discusses the use of sweep as a remedial strategy to control the aerodynamic limits in low-speed axial fan rotors. In this respect, the present work contributes to the understanding of the potential effect of blade lean on the shifting of the rotor stall margin. Numerical investigations have been undertaken on highly loaded fans of non-free vortex design, with the ideal total head rise coefficient typical of the industrial application range. Two rotors with identical nominal design parameters and, respectively, with 35° forward swept blades and unswept blades have been studied. The investigation has been carried out using an accurate in-house developed multilevel parallel finite element RANS solver, with the adoption of a non-isotropic two-equation turbulence closure. The pay-off derived from the sweep technology has been assessed with respect to the operating range improvement. To this end, the flow structure developing through the blade passages and downstream of the rotors, as well as loss distributions, have been analysed at design and near-peak pressure operating conditions. The analyses of three-dimensional flow structures showed that, sweeping forward the blade, the non-free vortex spanwise secondary flows are attenuated, and a control on the onset of stall is recovered. Moreover, the swept rotor features a reduced sensitivity to leakage flow effects. Consequently, it operates more efficiently approaching the throttling limit.
This paper describes the introduction of 3D blade designs into the core compressors for the Rolls-Royce Trent engine with particular emphasis on the use of sweep and dihedral in the rotor designs. It follows the development of the basic ideas in a university research project, through multistage low-speed model testing, to the application to high pressure engine compressors. An essential element of the project was the use of multistage CFD and some of the development of the method to allow the designs to take place is also discussed. The first part of the paper concentrates on the university-based research and the methods development. The second part describes additional low-speed multistage design and testing and the high-speed engine compressor design and test.