Heim_Ghandhi_JAutoEng_2011

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Engineering
Engineers, Part D: Journal of Automobile
Proceedings of the Institution of Mechanical
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The online version of this article can be found at:
DOI: 10.1177/0954407011404763
originally published online 15 June 2011
2011 225: 1067Proceedings of the Institution of Mechanical Engineers, Part D: Journal of Automobile Engineering
D M Heim and J B Ghandhi
Investigation of swirl meter performance
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Investigation of swirl meter performance
D M Heim and J B Ghandhi*
Engine Research Center, University of Wisconsin-Madison, Wisconsin, USA
The manuscript was received on 17 December 2010 and was accepted after revision for publication on 3 March 2011.
DOI: 10.1177/0954407011404763
Abstract: The performance of vane- and impulse-type swirl meters was investigated, and a
direct calibration method for swirl meters was developed. The zero-swirl bias of the meters
was tested by installing an axially aligned tube on the swirl meter. Both the vane- and
impulse-type meters showed insignificant zero-swirl bias. A known swirl was provided to the
swirl meters using an offset, inclined tube arrangement. The angular momentum flux deliv-
ered by this system was found to depend linearly on the product of the offset distance and
cosine of the inclination angle. Both the impulse- and vane-type meters were found to give
measurements below the known swirl value, but both meters gave results that were linearly
dependent on the angular momentum flux, which allows characterization of the meter’s effi-
ciency with a single parameter. The efficiency of the impulse-type meter varied from 0.7 to
0.93, was a moderate function of the flow straightener aspect ratio, and depended slightly on
the meter size. The vane-type meter’s efficiency was 0.32–0.45 for the conditions tested, was
insensitive to the paddle wheel flow straightener aspect ratio, and depended significantly on
the meter size. The vane-type meter measurements were also found to depend on the paddle-
to-bore-diameter ratio; values slightly exceeding unity should be used. The swirl meter effi-
ciency can be used to correct measurements to an absolute basis. Based on these findings, a
universal correction factor does not exist, and a given measuring device will need to be cali-
brated using the methodology described.
Keywords: swirl characterization, in-cylinder engine flow, port flow
1 INTRODUCTION
The combustion rate in an internal combustion
engine has long been understood to depend on the
in-cylinder mixture turbulence and the turbulence is,
in turn, directly influenced by large-scale flow struc-
tures such as swirl and tumble. It has been shown
that higher levels of swirl produce higher levels of
turbulence and lower cyclic variation [1]. Thus, the
level of swirl produced in an engine can directly
affect engine performance, cycle-to-cycle variability,
and emissions. Therefore, being able to measure
accurately and easily the swirl characteristics pro-
duced by an engine head is of great importance.
As early as 1934, Alcock [2] described using an in-
cylinder rotating vane to measure an optimal swirl
ratio that gave the best performance for a given
engine. These studies, however, required special
engine heads to accommodate placement of the vane
inside the cylinder and eliminate obstructions due to
poppet valves and injectors. Steady flow tests with a
vane-type meter subsequently replaced the in-cylin-
der rotating vane measurements, and have been used
for decades. Fitzgeorge and Allison [3] measured
swirl speed using a two-bladed impeller inside a flow
rig cylinder. They adjusted the axial distance between
the impeller and engine head and found the impeller
speed was a maximum when this distance was
1.4 times the cylinder bore diameter. They also used
the steady swirl results to try to predict the swirl in
an actual engine. Jones [4] measured swirl speed
using a straight-bladed anemometer inside the
flow rig cylinder and Watts and Scott [5] used a
*Corresponding author: Engine Research Center, University of
Wisconsin-Madison, 1500 W. Engineering Dr., Madison, WI
53706, USA.
email: ghandhi@engr.wisc.edu
1067
Proc. IMechE Vol. 225 Part D: J. Automobile Engineering

rectangular-shaped vane in their flow rig cylinder
and noted the form of the vane had little influence
on the measured swirl. Tindal and Williams [6] stud-
ied the air flow patterns in a steady-flow rig using
light paper flags and a vane anemometer to measure
swirl speed. They simulated the presence of a piston
in the cylinder by inserting a restrictor plate into the
flow rig liner at two bore diameters away from the
cylinder head and found that it caused the axial velo-
city to assume a more regular pattern, which resulted
in an increase in measured swirl.
Tippelmann [7] set forth the idea of using an
impulse-type swirl meter with a flow straightener
that converted the angular momentum into a mea-
surable torque. Uzkan et al. [8] described their
impulse-type meter having a honeycomb with small
cells and large aspect ratio capable of straightening
the swirling flow completely. They also note that the
honeycomb should not be inserted into the rig
cylinder (as in references [7] and [9]), but should lie
below it with a larger diameter to eliminate air
blow-by. Swirl measurements were made using dif-
ferent head-to-honeycomb distances. A monotonic
decrease in measured torque with increasing dis-
tance was observed and attributed to cylinder wall
friction. They estimated the rate at which the angu-
lar momentum decays is on the order of 10 per cent
per cylinder diameter of axial distance.
A number of studies have made comparisons
between vane-type and impulse-type meters and in
general conclude that vane-type meters provide
lower swirl coefficients than impulse-type meters.
Tippelmann [7] showed that the readings from a
vane-type anemometer were too small and varied in
magnitude compared with the impulse-type meter.
Monaghan and Pettifer [10] calculated swirl ratios
for four different types of ports using both vane-type
and impulse-type meters. Swirl ratios using the
impulse-type meter were generally 30 per cent great-
er than those using the vane-type meter. Stone and
Ladommatos [11] took cylinder head swirl measure-
ments using both a paddle wheel anemometer
and impulse-type meter and also concluded that the
paddle wheel results fell below those of the
impulse-type torque meter. Snauwaert and Sierens
[12] acquired steady rig swirl measurements with a
paddle wheel anemometer, impulse-type meter and
a laser Doppler velocimeter (LDV) to show that dif-
ferent flow patterns produced over the range of
intake valve lift have varying effects on measure-
ment accuracy. Tanabe et al. [13] tested the same
sized honeycomb using a vane wheel anemometer
and impulse-type meter. The vane wheel anem-
ometer gave swirl numbers below those of the
impulse-type meter, where the level of difference
depended on port type and valve lift. At the maxi-
mum valve lift the swirl numbers from the vane
wheel anemometer all tended to be about 0.8 times
as large as those calculated using the impulse-type
meter.
There have been limited published results on the
effect of flow straightener and paddle wheel geome-
try on measurement accuracy. In one such study,
Tanabe et al. [13] tested honeycomb flow straigh-
teners with cell sizes of 3.2 and 6.4 mm and heights
of 10, 20, and 30 mm. They first measured the drag
coefficients of the flow straighteners with steady
axial flow. Swirl numbers were then measured with
an impulse-type meter for three different cylinder
head port types at maximum valve lift. They found
that the flow straighteners with smaller drag coeffi-
cients tended to measure higher swirl numbers, but
the differences varied with valve lift.
Several investigators have made LDV measure-
ments to compare with steady-flow measurements.
Monaghan and Pettifer [10] took LDV measure-
ments in the steady-flow device to show how both
vane-type and impulse-type meters affect the axial
and radial velocity profiles in the swirl rig. The axial
flow was shown to be highly non-uniform and high-
er towards the outer part of the cylinder. This dis-
credited the assumption of uniform axial velocity
inherent in the use of the vane-type swirl meter cal-
culations, which leads to an underestimation of cal-
culated angular momentum. Kent et al. [14] made
LDV measurements in a motored engine, then inte-
grated the results to find the mean swirl at the end
of induction. The results were approximately 15 per
cent higher than predicted by the impulse-type swirl
meter, but they found their predictions of in-
cylinder swirl based on steady-flow angular momen-
tum flux measurements to be in trendwise agreement
with the LDV measurements in the motored engine.
In the current investigation, two geometrically
identical engine heads have been built to study the
speed- and size-scaling relationships of engine
flows. The length scale ratio between the engine
heads of this study is 1.69. Geometrically similar
engine heads should produce similar levels of swirl
when appropriately non-dimensionalized. In order
to span a wide range of in-cylinder conditions, the
heads are fitted with both normal and shrouded
intake valves. The first step in ensuring the flow
similarity of the heads was to perform steady-flow
measurements. These measurements, which span a
wide dynamic range in swirl level, brought to light
several features of steady-flow swirl measurements
that needed to be resolved in order to assess the
flow similarity between the scaled engine heads.
Vane- and impulse-type meters have been tested,
1068 D M Heim and J B Ghandhi

and an absolute calibration methodology has been
established in order to compare the results confi-
dently from the different sized heads.
2 EXPERIMENTAL APPARATUS
Testing was performed using a SuperFlow 600 flow
bench. The flow bench pulls air into the machine
with a prescribed pressure drop across an attached
test section ranging from 0.25 to 12 kPa (1 to 48 in
H2O). The volumetric flowrate was found from the
pressure drop across a calibrated orifice inside the
flow bench using an inclined manometer. The den-
sity of the air was calculated from temperature and
humidity data acquired using a Mannix model J411-
TH digital hygro thermometer, and the barometric
pressure was measured using a Heise model CM dial
pressure gauge. The engine heads were tested at the
industry-standard pressure drop of 7 kPa (28 in H2O).
The steady-flow swirl testing was performed
using a different swirl adapter fixture for each size
of engine head (see Fig. 1 and Table 1). Hereafter,
the two engine heads and associated components
will be referred to as ‘small’ for the smaller engine
and ‘large’ for the larger engine. The swirl adapter
fixtures, which have H/B = 1.5, are installed between
the cylinder head and the swirl meter. The valve lift
was adjusted using a modified micrometer that
mounted to the engine head. Intake horns, with
radii of curvature large enough to minimize the
pressure drop at the inlet to the intake ports, were
connected to the entrance of the intake ports. The
bore diameters of the swirl adapter fixtures were
the same as the engine cylinder bore. Table 1 gives
the relevant dimensions of the engine heads and
swirl adapter fixtures used with both vane-type and
impulse-type swirl meters.
The vane-type swirl meter used for this study was
an Audie Technology paddle-style swirl meter. The
meter featured a honeycomb paddle wheel made of
polycarbonate plastic with tubular cells. The outer
diameter of the paddle featured a smooth, thin poly-
carbonate sheet wrapped around the honeycomb to
form a continuous cylinder-like shape. The swirl
meter provided an electronic output of two pulses
per revolution, which are also used to determine
both the direction of rotation and the rotation rate
with the addition of an HP model 5315A universal
counter; for all testing, data were collected and aver-
aged over a 40 s period to obtain the mean rotation
rate of the paddle. The relevant dimensions of the
paddle meter are provided in Table 1.
Two impulse-type swirl meters were used for this
study. Tests were first conducted using the impulse-
type meter from the study by Bottom [15]. This will
be referred to as the ‘first’ impulse-type meter. In this
meter, a polycarbonate shaft is fixed on one end and
the other end is attached to a honeycomb flow
straightener consisting of an aluminium honeycomb
matrix (see Table 1 for dimensions). The shaft was
instrumented with two Vishay/Micro-Measurements
torsional strain gauges located 180° apart. The shaft
was designed to deform elastically for low angular
momentum flows. An Omega DMD-465 strain-gauge
amplifier provided the excitation voltage for the
strain gauges and a data acquisition system recorded
the instantaneous voltage at a rate of 10 Hz. Data
were collected and averaged over a 40 s period to
obtain the mean voltage (torque).
A second impulse-type meter was used for this
study, henceforth referred to as the ‘second’
Fig. 1 (a) Vane-type swirl meter test set-up; (b) impulse-type torque meter test set-up
Investigation of swirl meter performance 1069

impulse-type meter. In this meter, a Transducer
Techniques RTS-5 torque sensor was secured at the
bottom and, similarly to the first impulse-type meter,
a shaft was attached at one end to the sensor and on
the other end to a honeycomb flow straightener. The
honeycomb was made of the same material and tub-
ular structure as used in the vane-type meter. The
design of the second impulse-type meter allowed dif-
ferent honeycomb flow straighteners to be easily
tested. A Daytronic model 3270 strain gauge condi-
tioner/indicator provided the excitation voltage for
the torque sensor and the same data acquisition sys-
tem recorded the instantaneous voltage at a rate of
10 Hz. Dimensions of the honeycomb flow straigh-
tener will be discussed in a later section.
Both impulse-type meters were calibrated by
applying a set of known torques to the centre of the
honeycomb flow straightener. For each applied tor-
que, a corresponding voltage was recorded. Before
and after each applied torque, the zero-torque
voltage was recorded and averaged. The average
zero-torque voltage was subtracted from the applied-
torque voltage to obtain the voltage difference. The
voltage difference was plotted against the applied
torques to determine a linear calibration curve.
Calibration data were collected for counterclockwise
torques applied to the honeycomb flow straightener.
3 FLOW PARAMETERS
The flow parameters that will be used to character-
ize the engine heads are the flow coefficient, Cf, the
swirl coefficient, Cs, and the swirl ratio, Rs. The flow
and swirl coefficients are measured at discrete valve
lifts over the range of the cam profile, and are
reported as a function of L/D, where L is the valve
lift and D is the valve inner seat diameter. The flow
coefficient is a measure of the actual mass flowrate
to a theoretical mass flowrate and is defined as
Cf =
_m
rVBAv
(1)
where _m is the air mass flowrate, r is the upstream
air density, AV is the valve inner seat area, and VB is
the Bernoulli velocity given by
VB =
ffiffiffiffiffiffiffiffiffi
2DP
r
s
(2)
where DP is the pressure drop across the test sec-
tion. The incompressible relation for velocity is suf-
ficient at the 7 kPa pressure drop, i.e. the Mach
number is 0.32, but for higher pressure drops a
compressible form of the velocity should be used.
The swirl coefficient, Cs, is a characteristic non-
dimensional rotation rate and is calculated for vane-
type meters using
Cs =
vB
VB
(3)
where v is the vane or paddle wheel angular velocity
and B is the cylinder bore. For impulse-type swirl
meters, the swirl coefficient is calculated from
Cs =
8T
_mVBB
(4)
where T is the torque measured by the meter. The
swirl ratio, Rs, is a convenient single metric that
takes into account the flow and swirl coefficients
over the entire lift profile of the engine. The swirl
ratio is calculated as
Rs =
ph2
vBS
4AV
ÐuIVC
uIVO
CfCsdu
ÐuIVC
uIVO
Cfdu
!2 (5)
where hv is the volumetric efficiency, assumed equal
to 1 for all calculations, uIVO and uIVC are the crank
angle (rad), at intake valve open and intake valve
closed, respectively, and S is the engine stroke. A full
derivation of the swirl coefficient and swirl ratio can
be found in Appendix 2.
Table 1 Relevant dimensions of the engine heads,
swirl adapters, vane-type meter, and first
impulse-type meter
Parameter
Dimensions (mm)
Large engine head Small engine head
B 82.0 48.6
H 123.0 72.8
D 35.0 20.7
Lmax 7.9 4.7
DP 132.0
HP 15.9
DI 165.0
HI 64.0
Paddle honeycomb 3.7
cell diameter, dP
First impulse torque 6.4
honeycomb
cell size

4 INITIAL MEASUREMENTS
Initial swirl measurements on the geometrically
similar heads were performed using the first
impulse-type meter. The study included using both
standard and shrouded intake valves, where an 180°
shroud was used to produce higher levels of swirl.
The flowrate data indicated that the mass flowrate
was well scaled between the two heads, i.e. a certain
level of similarity had been achieved.
From equation (4) it can be seen that for a con-
stant pressure drop (VB) the torque will scale as the
swirl coefficient and a characteristic length to the
third power. Based on the 1.69 scale ratio, and
assuming a worst-case scenario of a swirl ratio of 3
for the large head with the shrouded valve and a
swirl ratio of 0.15 for the small head with a stan-
dard valve (a 20:1 ratio of Cs), a 96: 1 ratio of torque
is obtained. Thus, a measurement device with a
very high dynamic range is required to cover the
entire test range of interest. It was desired to have
a precision of 1 per cent (a 100: 1 signal-to-noise
(S/N) ratio) in the measurements. The S/N ratio
was calculated based on the variability in the mea-
sured torque over the 40 s integration period. For
both heads fitted with the shrouded valves, the pre-
cision criterion was achieved by L/D~0.07. For the
unshrouded valve cases, at the maximum valve lift
the criterion was just satisfied for the large head,
but for the small head the peak S/N ratio was 25: 1
over all L/D. These results motivated the investiga-
tion of a vane-type meter because, intrinsically, a
rotation rate is easier to measure with a wide
dynamic range.
Figure 2 shows the swirl coefficients for both
heads with the shrouded valves as a function of L/D
measured with both the first impulse-type meter
and the vane-type meter. This condition was chosen
because of the high S/N achieved with the first
impulse-type meter. The impulse meter results
show a good degree of similarity – the resulting swirl
ratios were 2.65 and 2.75 for the large and small
heads, respectively. In contrast, the vane-type meter
results showed two disturbing features. First, the
measurements for both heads differed from the
impulse meter results. Second, the results for the two
heads differed quite significantly from each other;
the swirl ratio was 0.57 for the small head and 1.13
for the large head. The former problem is an issue of
absolute accuracy, which will be discussed below,
but the latter is an issue of the operation of the vane-
type meter and is discussed here.
Owing to the difference in the diameters of the
two swirl adapter fixtures, it was thought that there
might be a difference in air frictional losses from the
paddle outside the cylinder bore (the same sized
paddle was used for both heads). The portion of the
paddle outside the cylinder would experience air
friction tending to retard the motion of the paddle,
which is consistent with the lower Cs measured for
the small head. In order to test the effect of air fric-
tional losses on the rotational speed of the paddle,
custom paddles were fabricated of the same honey-
comb material and geometry as the original paddle
wheel, but with a smaller paddle diameter, DP. For
both the small and large heads, the ratio of the pad-
dle diameter to the swirl adapter fixture, DP/B, was
set to 1.2. Figure 3 shows the results of the constant
DP/B tests for the same conditions as Fig. 2. It can
be seen that by controlling DP/B the differences
between the two vane-type meter measurements
have been eliminated, and it may be concluded that
self-similarity has been achieved. It is also possible
that the gap between the paddle and bore adapter
affected the friction, but this was not expressly
tested. There are, however, still differences in the
absolute value of swirl coefficient between the
impulse- and vane-type meter measurements.
5 ABSOLUTE CALIBRATION OF SWIRL METERS
The wide dynamic range required for these experi-
ments suggests that more than one swirl meter may
be needed. However, based on the initial measure-
ments it is clear that using a vane-type meter for the
low range and an impulse-type meter for the high
-1.0
-0.8
-0.6
-0.4
-0.2
0.0
0.2
Cs
0.250.200.150.100.050.00
L/D
Impulse Meter
Vane Meter
open symbol - small head
filled symbol - large head
Fig. 2 Initial measurements of swirl coefficient using
the first impulse-type meter and the vane-type
meters with the standard rotor. The measure-
ments are for the shrouded valves

range is not a good option unless an absolute refer-
ence can be established against which both meters
can be calibrated. There are two aspects to an abso-
lute calibration, establishing a zero point and deter-
mining the constant of proportionality (assuming a
linear dependence). Additionally, it is useful to moni-
tor the long-term performance of a flow bench, and
the apparatus that has been developed for calibration
can also be used for this.
5.1 Zero-swirl reference
A zero-swirl reference fixture is shown in Fig. 4(a)
with the relevant dimensions given in Table 2. The
zero-swirl reference features a tube that is coaxial
with the swirl adapter fixture and a flat plate that
secures to the top of the swirl adapter fixture. A flow
straightener was installed at the inlet of the tube in
order to help ensure a uniform incoming flow.
Tests were performed at flowrates corresponding
to a pressure drop of 28 in H2O across the test sec-
tion. The data from the impulse meter showed a
small torque offset in comparison to a zero-flow
condition. Converting this into an equivalent swirl
coefficient based on the measured flowrate, the
maximum value was Cs = 0.024, which is small in
comparison with typical values of Cs, even with the
unshrouded valve. The paddle meter results were
more difficult to quantify because the paddle was
essentially stationary, changing position erratically
but not rotating. It is sufficient to say that both swirl
meters were robust relative to a zero-swirl bias.
5.2 Known swirl reference
The second fixture is a known swirl reference and is
shown in Fig. 4(b) with the relevant dimensions giv-
en in Table 2. The known swirl reference features a
tube with its axis offset from the swirl adapter fix-
ture axis and a flat plate that secures to the top of
the swirl adapter fixture. The tube is installed in the
flat plate at an inclination angle uR relative to the
horizontal. Again, a flow straightener was installed
at the inlet of the tube.
For a given geometry (R2 and uR), it can be shown
(see Appendix 3) that the angled-tube geometry pro-
vides a constant value of Cs; the correlation for any
geometry is provided in Appendix 3. Thus, using
equation (4), T can be found, or combining equa-
tions (3) and (4), a measured v can be used to find
an equivalent torque, Teq, as a function of the mea-
sured velocity V, which is determined from the vol-
ume flowrate and pipe area and is used in place of
the Bernoulli velocity. In the subsequent plots, the
term angular momentum flux will be used, which is
equivalent to T (see equation (6) in Appendix 2).
Figure 5 shows the results of the angled-tube cali-
brations of the vane- and impulse-type swirl meters
for both the large and small fixtures (the ‘second’
impulse-type meter was used for these measure-
ments). For the vane meter measurements DP/
B = 1.2 was used, and for all cases the cell aspect
ratio (HI/dI or HP/dP) was 4.3. Both measurement
techniques show excellent linearity with respect to
the angular momentum flux, but there is not a 1: 1
correspondence between the measured (or derived
in the case of the vane meter) torque and the inlet
angular momentum flux. The high degree of linear-
ity indicates that a single conversion efficiency can
be used to describe the performance of the swirl
meters, and this efficiency is the slope of the lines in
Fig. 5. For the data in Fig. 5, the efficiency ranges
from 0.90 for the large fixture using the impulse-
type meter, to 0.32 for the small fixture using the
vane-type meter. From Fig. 5 it is clear that the con-
version efficiency is a function of the meter type
and the fixture size. The impulse-type meter gives
results that are larger in magnitude than the vane-
type meters by nearly a factor of two, and the
impulse-meter results are closer to, but still less
than, the correct value. For the data in Fig. 5, the
smaller fixture gave higher results for both meters.
The effect of the flow straightener or vane cell size
was measured using polycarbonate honeycombs
having a tubular geometry. The honeycomb cell dia-
meters tested were 6.4 and 3.7 mm. For the vane-type
meter DP/B was again set to 1.2 to minimize the fric-
tional losses, and the honeycomb height was limited
-1.0
-0.8
-0.6
-0.4
-0.2
0.0
0.2
Cs
0.250.200.150.100.050.00
L/D
Impulse Meter
Vane Meter
open symbol - small head
filled symbol - large head
Fig. 3 Swirl coefficient using the paddles with Dp/
B = 1.2 for the vane-type swirl meter. The
impulse-type meter measurements are the
same as Fig. 2

to 15.9mm (Hp/dp = 4.3) by the meter design. For the
impulse-type meter longer honeycombs were tested,
up to HI/dI = 17, and a fixed straightener diameter of
DI = 104 mm was used. The cell geometry results are
shown in Fig. 6. The vane-type meter (Fig. 6(a)),
showed a weak sensitivity to the cell geometry, but as
was seen in Fig. 5 the conversion efficiency is poor.
For the large fixture, the conversion efficiency was
~0.32 and for the small fixture it was ~0.44. The lower
conversion efficiency for the large fixture could be
due to friction at the hub, which would be greater for
the larger vane size, or from slip between the air and
the paddle. If air slip was causing the low conversion
efficiency, it might be expected that the higher HP/dP
cases would perform better, which was not the case.
The impulse-type meter showed a stronger sensi-
tivity to the flow straightener geometry, with the con-
version efficiency decreasing with increasing aspect
ratio of the honeycomb. This result agrees with the
findings of Tanabe et al. [13]. In comparison with the
vane-type meter, the conversion efficiency of the
impulse-type meter is significantly larger, but differ-
ences exist between the two fixture sizes and the
magnitude of the conversion efficiency can be as low
as 0.7. Thus, the results from an impulse-style meter
Fig. 4 Calibration devices for establishing (a) a zero-swirl reference and (b) a known swirl
condition
Fig. 5 Impulse- and vane-type meter responses to a
known angular momentum flux produced from
the angled tube for the small and large fixture.
For all cases a cell height-to-length ratio of 4.3
was used
Table 2 Dimensions of the vertical and angled refer-
ence standards
Parameter Dimensions
uR Vertical reference, 90° Angled reference, 45°
SR 127.0 mm
DR 19.0 mm
LR 445.0 mm

will underpredict the true level of swirl. It is possible
that the losses in the H/B = 1.5 cylinder diameter tube
could account for some of the underprediction seen
with the impulse-type meter.
6 RAW DATA CORRECTION
The swirl meter efficiency, such as those found in
Figs 5 and 6, can be used to correct measurements
to an absolute basis, and for the current study to
remove size-dependent measurement artefacts. The
correction procedure simply involves dividing the
measured torque by the efficiency factor deter-
mined using the angled-tube measurements.
Figures 7(a) and (b) show both uncorrected and
corrected data for the vane- and impulse-type swirl
meters acquired using the large head. The data are
shown using the same axis range to highlight the
dynamic range of the measurements. The vane-type
measurements were made with Hp/dp = 4.3 and the
impulse-meter measurements were made with HI/
dI = 1.4. Similar to the results of Fig. 3, the uncor-
rected vane-type meter results are approximately a
factor of two lower in magnitude for the shrouded
valve case (Fig. 7(a)). After correction, this difference
is significantly reduced, and the vane-type meter
measurements slightly exceed the impulse-type
measurements. The unshrouded valve data, which
exhibit very low values of Cs, are slightly overcor-
rected, but since the swirl level is not very signifi-
cant, this is not too problematic.
7 CONCLUSIONS
A methodology to measure the absolute perfor-
mance of swirl meters was developed. An axial tube
Fig. 6 Swirl conversion efficiency as a function of the cell aspect ratio for (a) the vane-type meter
and (b) the impulse-type meter
Fig. 7 Raw and corrected swirl coefficient data for (a) a shrouded valve and (b) an unshrouded
valve and the large head

arrangement was employed to determine the zero-
swirl performance of a meter. Both the vane- and
impulse-type meters tested showed insignificant
zero-swirl bias. An offset, angled-tube arrangement
was developed to measure swirl meter performance
against a known swirl reference. The swirl coeffi-
cient for the angled-tube geometry was found to be
a linear function of the product of the offset dis-
tance and the cosine of the inclination angle.
Impulse-type swirl meters were found to give mea-
sured results closer in magnitude to the known swirl
level than vane-type meters; both meter types
showed a linear dependence on the input angular
momentum flux, allowing calibration using a single
coefficient. The absolute efficiency of the impulse-
type meter was found to be a function of its physical
size, and of the geometry of the flow straightener
used, with lower cell aspect ratios giving higher effi-
ciency. Vane-type meters were found to be sensitive
to the paddle-to-bore-diameter ratio; higher values
of Dp/B give lower measured swirl coefficient due to
excess friction. The efficiency of the vane-type
meter was found to be insensitive to the paddle cell
aspect ratio, but was sensitive to the physical size of
the meter, even with a constant Dp/B.
ACKNOWLEDGEMENTS
Support for this work was provided by the Wisconsin
Small Engine Consortium. The authors’ special
thanks are extended to D. Kilian for designing the
new impulse-type meter and for his help in data
collection.
Ó Authors 2011
REFERENCES
1 Bracco, F. Structure of flames in premixed-charge
IC engines. Combust. Sci. Technol., 1988, 58, 209–
230.
2 Alcock, J. Air swirl in oil engines. Proc. Instn Mech.
Engrs, 1934, 128, 123–193.
3 Fitzgeorge, D. and Allison, J. Air swirl in a road-
vehicle diesel engine. Proc. Instn Mech. Engrs,
1962–1963, 4, 151–177.
4 Jones, P. Induction system development for high-
performance direct-injection engines. Proc. Instn
Mech. Engrs, 1965–1966, 180(Part 3N), 42–52.
5 Watts, R. and Scott, W. Air motion and fuel distri-
bution requirements in high-speed direct injection
diesel engines. Proc. Instn Mech. Engrs, 1969–1970,
184(Part 3J), 181–191.
6 Tindal, M. and Williams, T. An investigation of
cylinder gas motion in the direct injection diesel
engine. SAE paper 770405, 1977.
7 Tippelmann, G. A new method of investigation of
swirl ports. SAE paper 770404, 1977.
8 Uzkan, T., Borgnakke, C., and Morel, T. Charac-
terization of flow produced by a high-swirl inlet
port. SAE paper 830266, 1983.
9 Davis, G. and Kent, J. Comparison of model calcu-
lations and experimental measurements of the bulk
cylinder flow processes in a motored PROCO
engine. SAE paper 790290, 1979.
10 Monaghan, M. and Pettifer, H. Air motion and its
effect on diesel performance and emissions. SAE
paper 810255, 1981.
11 Stone, C. and Ladommatos, N. The measurement
and analysis of swirl in steady flow. SAE paper
921642, 1992.
12 Snauwaert, P. and Sierens, R. Experimental study
of the swirl motion in direct injection diesel
engines under steady-state flow conditions (by
LDA). SAE paper 860026, 1986.
13 Tanabe, S., Iwata, H., and Kashiwada, Y. On char-
acteristics of impulse swirl meter. Trans. Jap. Soc.
Mech. Engrs, Ser. B, 1994, 60(571), 1054–1060.
14 Kent, J., Haghgooie, M., Mikulec, A., Davis, G.,
and Tabaczynski, R. Effects of intake port design
and valve lift on in-cylinder flow and burnrate. SAE
paper 872153, 1987.
15 Bottom, K. PIV measurements of in-cylinder flow
and correlation with engine performance. PhD The-
sis, University of Wisconsin–Madison, Wisconsin,
USA, 2003.
APPENDIX 1
Notation
Av valve inner seat area
B swirl adapter fixture bore
Cf flow coefficient
Cs swirl coefficient
dI diameter of impulse torque meter
honeycomb cells
dP diameter of paddle meter honeycomb
cells
D inner seat diameter
DI diameter of impulse torque meter
honeycomb flow rectifier
DP diameter of paddle meter paddle wheel
DR diameter of reference standard tubes
H height of swirl adapter fixture
HI height of impulse torque meter honey-
comb flow rectifier
HP height of paddle meter paddle wheel
L valve lift
Lmax peak valve lift
LR calibration tube length
_m mass flowrate of air
P pressure

Rs swirl ratio
R1 radius of angled reference standard tube
R2 angled reference standard offset from the
centre-line of the cylinder bore
S engine piston stroke
S/N signal-to-noise ratio
SR flow straightener length
T measured torque
Teq equivalent torque
V measured velocity
VB Bernoulli velocity
hv volumetric efficiency
uIVC crank angle at intake valve closed
uIVO crank angle at intake valve open
uR angle of reference tube
r density of air
v paddle wheel angular velocity
APPENDIX 2
Swirl coefficient and swirl ratio determination
Consider a cylinder closed at one end and open at
the other end. Fluid enters the cylinder through
some arbitrary surface S1 on the closed end, and
flows uniformly (in the axial direction) out of the
open end of the cylinder. If the exit plane contains a
flow-straightening device such that the exit flow is
purely in the axial direction, then the torque
required to hold the flow straightener is found from
the conservation of angular momentum as
Tz =
ð ð
S1
rvurv Á dA
(6)
If instead, the flow exits the open end of the cylin-
der with a solid-body rotation at rotational rate v,
then using angular momentum conservation the
rotation rate can be written in terms of Tz from
equation (6) as
v =
8Tz
_mB2
(7)
The swirl coefficient is defined as v normalized by a
characteristic rotation rate VB/B, which using equa-
tion (7) gives
Cs =
8Tz
_mVBB
(8)
The swirl ratio, Rs, is found by considering the
unsteady angular momentum conservation for the
cylinder
0 =
∂
∂t
ð ð ð
8
r3vð Þrd8
2
4
3
5 +
ð ð
S1
rvurv Á dA
(9)
The second term on the right-hand side is just Tz
using equation (6), and by assuming a quasi-steady
filling process for an initially empty cylinder, equa-
tion (9) may be integrated to find
ð ð ð
8
r3vð Þrd8 =
ðtIVC
0
Tzdt (10)
Assuming that the cylinder contents have a solid-
body rotation at O at the time of intake valve clo-
sure, and that the engine rotation rate is Oeng, then
the swirl ratio is found as
Rs[
O
Oeng
=
32
rpSB4O2
eng
ðuIVC
uIVO
Tzdu (11)
Substituting for Tz from (8) and collecting in terms
of Cf, the following is found
Rs =
4V 2
B Aref
pSB3O2
eng
ðuIVC
uIVO
CfCsdu (12)
Where Aref is the reference area used to define Cf.
In order to remove the engine speed from the
denominator of equation (12), the following is noted
ðuIVC
uIVO
Cfdu =
Oeng
rVBAref
ðtIVC
0
_mdt (13)
and that the rightmost integral is just the mass
delivered per cycle. Assuming that the swept volume
is close to the total cylinder volume, and using the
volumetric efficiency, hv, the following is found
Oeng =
4VBAref
hvpB2S
ðuIVC
uIVO
Cfdu (14)
Introducing equation (14) into equation (12), the
following is found
Rs =
h2
vpSB
4Aref
ÐuIVC
uIVO
CfCsdu
ÐuIVC
uIVO
Cfdu
2 (15)
The results presented herein used the inner seat
area for Aref, but inspection of equation (15) sug-
gests that if Aref were chosen to be the cylinder

cross-sectional area pB2
=4 further simplifications
would be achieved.
APPENDIX 3
Known swirl calibration
The torque measured by a flow straightener can be
written as
Tz =
ð ð
S1
rvurvndA
(16)
where the normal velocity, vn = V cos u, the tangen-
tial velocity, vu = V cos u cos a, a is the angle that
the differential area element dA makes with the ver-
tical in Fig. 4(b), and V is the measured velocity
obtained from the flowrate measurement of the flow
bench and the known pipe area. Normalizing all of
the dimensions by the cylinder radius (B/2) and
denoting dimensionless distances with an overbar,
e.g. R1 = R1= B=2ð Þ, and recasting equation (16) in
terms of the swirl coefficient, the following is found
Cs =
4
p R2
1
sin u cos u
ð ð
S1
r cos a d A (17)
The integral was evaluated numerically, and the
results were found to be independent of tube size
R1, linearly dependent on tube offset R2, and depen-
dent on the cos u. The results can be summarized in
the single plot shown in Fig. 8, where a single line
that passes through the origin fits all of the data
with a slope as shown.
These results have been used in conjunction with
equation (8) to find the torque as a function of vol-
ume flowrate, where V is used in place of VB.
Similarly, using equations (7) and (8), it is possible
to find the vane rotation rate as a function of vol-
ume flowrate.
Fig. 8 Swirl coefficient dependence on angled tube
geometry

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