-------
This initial increase of heat release rate is attributed to more rapid mixing of
the vapor fraction. However, insufficient initial penetration of fuel droplets
leads to an extended flame as indicated by the radial temperature gradients which
remain steeper for a greater axial distance. For the case of 0.3 swirl, the greater
part of the available chemical energy was released within an axial distance of
approximately six combustor diameters (X/D = 6) and the downstream temperature dis-
tributions are relatively flat. Isotherms corresponding to S = 0.6, on the other
hand, indicate a larger flame length, as evidenced by peak temperatures extending
the full length of the combustor (X/D = 14).
Mean axial gas- and droplet-velocity contours obtained for iso-octane/air
combustion at one atmosphere pressure are presented in Figs. 6 and 7. Differences
between the local gas and droplet velocities are apparent near the injector, as
are areas of flow recirculation. Also, some unsteadiness of the flow was indi-
cated by fluctuations in the droplet velocity measurements. Farther downstream
(X/D>1.0), droplet sizes and concentrations are reduced by evaporation and combus-
tion and droplets are convected at the local gas velocity. At S = 0.3, it was not
possible to distinguish between local fuel droplet and gas velocities in the
vicinity of the spray near the centerline because of the high droplet concentration.
Consequently, no gas flow recirculation is shown; however, the existence of a re-
circulation zone may be inferred from the droplet velocity data, Fig. 7a. Gas
velocity measurements were possible at S = 0.6 and a torroidal-shaped recirculation
zone was identified (Fig. 6) . A primary effect of increasing the inlet air swirl
from 0.3 to 0.6 was to shift the regions of droplet recirculation closer to the
injector, thereby influencing flame stabilization and energy release in the initial
region of the combustor.
The local time-mean axial velocities are somewhat higher at S = 0.6 and the
diffusion-flame-like flow field structure persists for a greater axial distance.
The insensitivity of the fuel droplet axial velocities to the level of inlet air
swirl is also apparent from the similar appearance of the droplet velocity distri-
butions in the initial mixing regions; however, at S = 0.6, high droplet velocities
persist farther downstream. Spray trajectories, as determined from laser
velocimetry and laser holography measurements are also shown in Fig. 7 and are in
good agreement with the nominal spray angle of 60 degrees. The persistence of
droplet velocity and the higher gas velocities at S = 0.6 are associated with the
increase in recirculation zone size and consequent increase in mass flux density
outside the recirculation zone. This increase in mass flux density was enhanced
by the confined geometry of the combustion facility.
Gas composition contours for low (S = 0.3), and moderately (S = 0.6) swirling
flows, discussed above, are shown in Figs. 8, 9 and 10. These data indicate the
tendency of increased swirl to suppress mixing beyond a zone of rapid initial
mixing near the injector. Radial concentration gradients are sharper with increased
swirl and CO burnout is slower. Further insight into the effect of swirl on
19
-------
FIG. 6
MEAN AND RMS GAS VELOCITY DISTRIBUTIONS
ISO-OCTANE/AIR, 1 ATM, TA|R - 533°K. 0 - 0.65
mm RECIRCULATION ZONE
URMS>10M/SEC
(A) SWIRL=0.3
-1.0
:
EC
a.
-
:
1
.
z.
i
oc
.25 .50
, AXIAL DISTANCE, X/D
.75
(B)SWIRL=0.6
-1.0
o
DC
c -0.5
<
c.
:
< 0.5
cc
'
SPRAY TRAJECTORY
- ^ _ < 20 M/S u = 20 M/SEC
40-
.50
AXIAL DISTANCE, X/D
20
-------
MEAN AND RMSDROPLE T VELOCITY DISTRIBUTIONS
ISO OCTANE/AIR. 1ATM, TA|H 533°K,
-------
(A) SWIRL=0.3
i.o r
0.5
O
tr
O
0.
Q
<
QC
-0.5
-1.0 L
(B) SWIRL=0.6
TIME-AVERAGED O2 DISTRIBUTIONS
ISO-OCTANE/AIR , 1 ATM, TA,R = 533°K , 0 = 0.65
15
10
3
5
10
6 8 10
AXIAL DISTANCE, X/D
12
FIG. 8
1
15
I I I I I I I
14
16
O
CE
cc
O
O-
Q
<
tr
0.5
55 0
-0.5
-1.0 L
14
6
2
1
1
2
6
14
1 I 1
1 I 1
6 8 10
AXIAL DISTANCE, X/D
12
14
16
22
-------
o
CC
o
a.
_l
<
Q
OC
TIME-AVERAGED CO DISTRIBUTIONS
ISO-OCTANE/AIR, 1 ATM, TA|R = 533°K, 0=0.65
FIG. 9
(A) SWIRL=0.3
1.0 i-
0.5
-0.5
-1.0
6810
AXIAL DISTANCE, X/D
12
14
16
O
tr
cc
O
h;
w
o
Q.
_J
<
Q
tE
(B)SWIRL=0.6
0.5
-0.5
-1.0
|
I
6 8 10
AXIAL DISTANCE, X/D
12
14
16
23
-------
FIG.10
(A) SWIRL=0.3
1.0r
0.5
o
cc
to
O
Q.
_J
<
Q
-0.5
-1.0
TIME-AVERAGED CO2 DISTRIBUTIONS
ISO-OCTANE/AIR, 1 ATM, TA,R = 533°K 0=0.65
I
1
6 8 10
AXIAL DISTANCE, X/D
12
14
16
(B) SWIRL=0.6
0.5
O
QC
IT
O
Q.
Q
<
CC
-0.5
-8
-7
5
-3
6 8 10
AXIAL DISTANCE, X/D
24
12
14
16
-------
iso-octane vaporization and combustion is obtained from comparisons of total
hydrocarbon concentration distributions (Fig. 11) together with profiles of the
percentage of hydrocarbons vaporized (Fig. 12). In the upstream section of the
combustor, total hydrocarbon concentrations are higher for the lower swirl number
(Fig. 11), while the concentrations of unvaporized fuel (computed from the product
of total hydrocarbon concentration from Fig. 11 and 1 minus the fraction vaporized
from Fig. 12) are about the same. In this upstream region increased swirl promotes
mixing of vaporized fuel. However, for the higher swirl number unvaporized fuel
persists further downstream and the total hydrocarbon concentrations are greater
in the downstream sections even after the fuel has vaporized. It is likely that
this is a result of reduced droplet penetration into the airstream with increased
swirl and the rapid decay of swirl (shear) induced mixing with axial distance.
NO concentration distributions are shown in Fig. 13. The regions of high NO
concentration within the combustor are coincident with the regions of locally
high temperature. At low swirl, higher concentrations of NO were measured close
to the combustor centerline and in the vicinity of the injector. In contrast,
at moderate swirl the reaction zone is moved rapidly outward and closer to the
injector, and NO formation occurs in a narrow annular region. The combined effect
of low oxygen concentration and low temperature result in a reduced rate of NO
formation and, therefore, lower NO exhaust emissions levels at S = 0.6.
The mapping data indicate that, in the present combustor configuration, in-
creasing swirl from 0.3 to 0.6 increases mixing of partially vaporized fuel with air
in the initial region of the flow, resulting in increased energy release rates.
Hence, increased swirl tends to move the region of flame stabilization closer to
the fuel injector. Beyond this initial region, increasing swirl appears to suppress
vaporization of the liquid fuel and subsequent mixing of the vaporized fuel with
air. At sufficiently high swirl numbers, the radial pressure gradients reduce
penetration of partially vaporized droplets into the airstream, resulting in a
relatively cold fuel-rich region on the combustor centerline. The reduced vapori-
zation rates result in an extended mixing region. In addition, axial decay of
swirl-induced shear levels tends to reduce mixing rates downstream from the injec-
tor. Reduced mixing rates result in generally slower fuel oxidation and CO burnout
rates and in lower NO formation rates.
Effect of Pressure On Flow Field Structure and NO Formation
Previous tests (Ref. 9) with natural gas have demonstrated several effects of
pressure on exhaust emissions at constant mass flow rate. Increasing pressure
from 1 to 3 atmospheres decreased CO levels and increased exhaust NO levels
principally because combustor residence time increases with pressure. However,
the local rate of energy release decreased indicating that mixing rate was
suppressed at higher pressure. This would be expected since shear levels decrease
25
-------
TIME-AVERAGED DISTRIBUTIONS OF UNBURNED HYDROCARBONS
ISO-OCTANE/AIR,1ATM,TA|R=533°K,$ =0.65
FIG.11
(A) SWIRL=0.3
1.0
0.5
cc
H
O
to
O
Q_
-0.5
DC
-1.0
0
_L
_L
2345
AXIAL DISTANCE,X/D
(B) SWIRL=0.6
O
a:
cc
O
\^_
CO
O
o.
_i
<
Q
1.0
0.5
-0.5
-1.0
2345
AXIAL DISTANCE,X/D
26
-------
FIG.12
TIME-AVERAGED PERCENTAGE OF HYDROCARBONS VAPORIZED
ISO-OCTAN E/AI R,1 ATM,TA| R=533°K.$ =0.65
(A) SWIRL=0.3
-1.0
_J
7
AXIAL DISTANCE, X/D
(B)SWIRL=0.6
O
DC
DC
O
CO
O
Q.
_l
<
O
cc
-0.5h
-1.0
3456
AXIAL DISTANCE, X/D
27
-------
FIG.13
TIME-AVERAGED NO DISTRIBUTIONS
ISO-OCTANE/AIR , 1 ATM, TA|R = 533°K ,(/>=0.65
(A) SWIRL=0.3
1.0
o
cc
cc
O
in
O
CL.
O
<
cc
0.5
-0.5
-1.0 I
.10
6810
AXIAL DISTANCE, X/D
12
14
(B) SWIRL=0.6
1.0 I-
o
cc
cc
z"
g
H
CO
O
a.
_i
<
Q
<
CC
0.5
-0.5
-1.01-
100
10
1
50
I
6810
AXIAL DISTANCE, X/D
12
14
16
28
-------
with decreasing velocity. This conclusion is supported by the fact that at even
higher pressure, 7 atm, exhaust hydrocarbon levels increased despite the increased
residence time.
For two of the liquid fuels investigated, No. 2 distillate oil and propane,
increasing pressure from 1 to 3.3 atmospheres decreased CO levels and increased
NO levels. However, when iso-octane was used as a fuel, CO levels increased and
NO emission decreased as the pressure was increased from 1 to 3.3 atmospheres.
The flow field mapping tests conducted using iso-octane confirm the results of the
input-output tests and indicate the effect of pressure on the flow field structure
and pollutant formation.
The time-mean temperature distributions, Fig, 14, indicate that longer flames
are obtained at higher pressure. Furthermore, peak temperatures are lower at higher
pressure suggesting lower energy release rates. Examination of the species concen-
tration distributions obtained for iso-octane/air combustion at elevated pressure
are consistent with the temperature data. Examination of Fig. 15 reveals that at
3.3 atm pressure, the 02 cencentrations near the combustor centerline are lower,
indicating a reduced mixing rate. Similarly, initial breakdown of the fuel to CO
and oxidation of CO to C02 is slower at 3.3 atm (cf., Figs. 16 and 17). Figure 18
shows that NO is formed in an annular region close to the injector at approximately
the same radial location as the peak temperature. There are steep radial gradients
and low NO concentration levels at the combustor centerline. Peak NO concentra-
tions at the elevated pressure are much lower than were observed at atmospheric
pressure. An increase in pressure from 1.0 atm to 3.3 atm results in a significant
decrease in NO emissions which may be attributed in part to lower temperatures.
One possible explanation for the different effect of pressure on flow field
structure for iso-octane in comparison with natural gas, propane and No. 2 dis-
tillate oil is as follows: The propane rapidly vaporizes on injection into the
combustor. Hence, both natural gas and propane may be considered gaseous fuels.
In spite of reduced mixing rates resulting from the reduced shear levels associated
with the lower air velocities, combustion is enhanced and NO emissions increase
due to increased residence time and increased reaction rates. In contrast, No. 2
distillate oil burns largely inhomogeneously since vaporization rates are rela-
tively low due to higher boiling points. Increased droplet penetration at higher
pressure partially offsets the effect of reduced mixing due to shear and combus-
tion goes to completion because of increased residence time. Iso-octane is more
volatile than No. 2 distillate oil, and droplet vaporization tends to limit drop-
let penetration. With relatively little penetration of iso-octane liquid, com-
bustion efficiency would be governed largely by droplet vaporization rates. But
the droplet vaporization rate is a function of droplet boundary layer thickness,
which in turn is a function of the product of gas density and relative velocity
between droplets and air. For air moving at a velocity higher than the droplet
velocity, as pressure increases the relative velocity might well be reduced far
29
-------
FIG. 14
TIME-AVERAGED TEMPERATURE DISTRIBUTIONS
ISO-OCTANE/AIR, TA)R = 533 °K,0= 0.65
(A) SWIRL = 0.3, 1 ATM
1.0r
0.5
DC
z"
O
2
_
< -0.5
Q
-1.0
4 6 8 10 12
AXIAL DISTANCE, X/D
14
16
(B) SWIRL = 0.3, 3.3 ATM
O
EC
OC
1.0
0.5
O
t 0
CO
O
Cu
< -0.5
Q
-1.0
4 6 8 10 12
AXIAL DISTANCE, X/D
14 16
30
-------
TIME-AVERAGED O2 DISTRIBUTIONS
ISO-OCTANE/AIR/I ATM,TA|R =533°K,0=0.65
(A) SWIRL=0.3. 1 ATM
1.0 r
0.5
O
tr
2
O
H
W3
O
Q.
_l
<
O
<
oc
-0.5
-1.0
'15
10
3
5
10
-15
6 8 10
AXIAL DISTANCE, X/D
12
FIG.15
14 16
(B) SWIRL=0.3, 3.3 ATM
1.0
0.5
O
or
ce
LU
0
2
<
1- 0
to
Q
_l
<
Q
<
EC
-0.5
-1.0 L
19
10
19
. _l J 1 '
6 8 10
AXIAL DISTANCE, X/D
12
14 16
31
-------
FIG. 16
TIME-AVERAGED CO DISTRIBUTIONS
ISO-OCTANE/AIR, TA)R = 533°K, 0= 0.65
(A) SWIRL = 0.3, 1 ATM
1.01-
o
cc
cc
CO
o
Q_
Q
<
QC
0.5 -
-0.5
-1.0
EXHAUST CO = 0.22%
6 8 10
AXIAL DISTANCE, X/D
12
14
16
(B) SWIRL = 0.3, 3.3 ATM
1.0r
O
cc
cc
CO
O
Q.
Q
<
CC
0.5
-0.5
-1.0
_L
EXHAUST CO = 0.25%
_L
J_
6 8 10
AXIAL DISTANCE, X/D
32
J
12
14 16
79-01-78-2
-------
TIME-AVERAGED CO2 DISTRIBUTIONS
ISO-OCTANE/AIR, ,TA|R = 533OKr <£ = 0.65
(A) SWIRL = 0.3, 1 ATM
1.0
O
or
O
O
Q.
0.5
-0.5
-i.oL
EXHAUST C02 = 8.15%
I I
I
6 8 10
AXIAL DISTANCE, X/D
FIG. 17
J I
12 14 16
(B) SWIRL = 0.3, 3.3 ATM
i.o
O
£T
0.5
co 0
O
o.
_j
<
Q
-0.5
-1.QL-
EXHAUST CO2 = 7.5%
J I J= J
6 8 10
AXIAL DISTANCE, X/D
33
12 1'
J
16
79-01-78-1
-------
FIG. 18
o
oc
10
O
a.
Q
<
DC
TIME-AVERAGED NO DISTRIBUTIONS
ISO-OCTANE/AIR , TAm - 533°K , 0= 0.65
(A) SWIRL=0.3,1ATM
1.0 r
0.5 -
-0.5
-1.0
1 1 1 1 1
1 1 1
6810
AXIAL DISTANCE, X/D
12
14
16
O
en
en
o
Q.
_l
<
Q
(B) SWIRL=0.3,3.3 ATM
1.0r
0.5
-0.5
1
3 2
1
4
l liii,
8 10 n >»
AXIAL DISTANCE, X/D
14
16
34
-------
more than the density is increased. This could reduce vaporization rate with
increasing pressure. Increasing boiling points with higher pressure would also
tend to reduce the heat transfer rate with vaporizing fuel which is proportional
to the difference between the ambient temperature and the boiling point temperature.
Increased reaction rates due to increased pressure and increased residence time
do not compensate for reduced shear levels, poorer penetration and lower vaporiza-
tion rates and a significant amount of fuel vaporized prior to burning.
Effect of Fuel Type On Flow Field Structure and NO Formation
Liquid propane, iso-octane and No. 2 fuel oil differ widely in the physical
properties which influence the atomization and vaporization (velocity, surface
tension, heat capacity, latent heat of vaporization, vapor pressure). The energy
added to the airstream by combustion at a given equivalence ratio is of similar
magnitude for each of the three fuels; thus little difference in flow field struc-
ture or emission levels can be expected on the basis of equilibrium thermodynamic
considerations. Also, the amount of fuel-bound nitrogen found in all of the fuels
is quite small, and thus this factor is not believed to contribute significantly
to the overall level of nitric oxide production. Typical properties of the liquid
fuels and the results of limited quantitative fuel analyses are given in Appendix E.
Significant differences existed between the temperature patterns observed in
the burner when using liquid propane as compared to patterns produced when using
iso-octane or fuel oil see Fig. 19. This difference in pattern is due largely
to the difference in the fuel distributions achieved when injecting propane. These
fuel pattern differences are illustrated in Fig. 20 which presents levels of total
unburned hydrocarbons within the combustor as determined by use of the phase-dis-
criminating probe. Most of the propane was found to be concentrated near the
centerline of the combustor; this fuel distribution is believed to have resulted
from flashing of the liquid propane within the injector with the result that a
conical spray was not achieved. Because of the initial fuel distribution, combus-
tion was slow and peak temperatures were not achieved in the initial regions of
the combustor (Fig. 19a) . In the case of the iso-octane and fuel oil, spray
patterns were similar and fuel penetrated to the outer combustor radii within two
test section diameters (see section on Spray Characteristics). Temperature patterns
produced were also qualitatively similar (Figs. 19b and 19c) , the most significant
difference being the higher temperatures at the outer radii of the combustor in
the case of the fuel oil. Combustion appears to be more intense in the case of
the iso-octane spray resulitng in slightly higher peak temperatures and steeper
temperature gradients. The lower volumetric heat release rates in the case of the
fuel oil are probably associated with the fuel oil droplet characteristics. The
fuel oil droplets were somewhat larger than the iso-octane droplets initially and
the fuel oil vaporizes less rapidly than iso-octane. Thus, although the distri-
bution of unburned fuel in the initial region of the combustor is qualitatively
35
-------
TIME-AVERAGED TEMPERATURE DISTRIBUTIONS
S =0.3,1 ATM,TA|R=533°K, 0=0.65
FIG. 19
(A) PROPANE
1.0
0.5
O
DC
o
H
w
o
Q.
S-0.5
DC
-1.0
4 6 8 10 12
AXIAL DISTANCE, X/D
14
-------
TIME-AVERAGED DISTRIBUTIONS OF TOTAL UNBURNED HYDROCARBONS
5 = 0.3,1 ATM, TA|R=533°K,<£=0.65
FIG. 20
(A) PROPANE
1.0
O
SE
CC
zf
O
O
CL
cc
-1.0L
(B) ISO-OCTANE
1.0
-0.5 -
i i I I I 1
-1.0
(C) NO.2 FUEL OIL
1.0
O
cc
cc
to
O
CL
-I
<
Q
CC
-0.5 -
23456
AXIAL DISTANCE,X/D
37
-------
similar for the fuel oil and the iso-octane (Figs. 20b and 20c), the fuel oil
droplets were larger and required greater time, and hence, distance to burn
completely. The fact that a greater amount of reaction took place in the outer
radii in the case of the fuel oil is confirmed by measurement of the oxygen con-
centration which shows that lower oxygen concentrations were found in this region
(Fig. 21).
Effect of Air Preheat On J?iow Field Structure and NO Formation
An increase in the inlet air temperature will influence flow field character-
istics by affecting flow velocities, chemical reaction rates, and heat transfer
rates. With all other conditions held constant, an increase in temperature will
result in correspondingly higher temperatures throughout the combustor and will
create higher flow velocities. These higher flow velocities will have the primary
effect of decreasing the residence time of the combustor gases. The diminished
time available for completion of the chemical kinetic processes is in most cases
more than offset by the strong temperature dependence of individual reaction rates.
Decreased time available for the droplet vaporization will be compensated for by
the increased heat transfer rate associated with the greater temperature difference
between the gas and the droplets.
Examination of the temperature patterns (Fig. 22) obtained for the case where
the entrance temperature was increased by 220K (a 40 percent increase) shows that
the combined effect is primarily to increase the temperature levels qualitatively,
the temperature pattern did not change significantly. Correspondingly, the inlet
temperature change resulted in only small changes in the unburned fuel pattern
(Figs. 23a and 23b). As would be expected, the fraction of the unburned fuel
existing in the vapor state was greater for the increased temperature level case
(Figs. 23c and 23d).
The rate of formation of nitric oxide is very sensitive to local temperature
and accordingly, the increased temperature levels resulted in an approximate
doubling of the local NO concentration ratios (Fig. 24). This dramatic increase
occurred over the complete equivalence ratio range tested in the input-output
experiments. Emissions of CO would be expected to decrease with increased preheat
level because of the increased rate of CO oxidation and higher temperature levels,
and this, indeed, was found to be the case.
In conclusion, the effect of the increased preheat level was primarily to
increase the temperature levels throughout the combustor and thereby to increase
the production of nitric oxide; temperature patterns and composition patterns
remain relatively unchanged.
38
-------
FIG. 21
TIME-AVERAGED O2 DISTRIBUTIONS
5=0.3/1
(A) ISO-OCTANE
1.0 |
O
cc
cc
0.5
15
10
to
O
c_
_l
<
Q
-0.5
-1.0 L
3
5
10
15
_L
J_
_L
6 8 10
AXIAL DISTANCE, X/D
12
14
16
(B) NO.2 FUEL OIL
1.0r-
0.5
O
cc
z
O
O
a.
Q
<
oc
-0.5
-1.0
6 8 10
AXIAL DISTANCE, X/D
12
14
16
39
-------
FIG. 22
TIME-AVERAGED TEMPERATURE DISTRIBUTIONS
N0.2 FUEL Ol L/AIR, 5 = 0.3 . 1 ATM, $ = 0.65
(A)TA|R = 533°K
1.0r
O
DC
in
O
Q_
oc
0.5
_
< -0.5
Q
-1.0L
1100
4 6 8 10 12
AXIAL DISTANCE, X/D
14 16
(B)TAm=750 °K
O
ec
CO
O
Q_
_l
<
Q
QC
1.0
0.5
-1.0
46 8 10 12
AXIAL DISTANCE, X/D
14
16
40
-------
TIME-AVERAGED DISTRIBUTIONS OF UNBURNED HYDROCARBONS
-NO.2 FUEL OIL/AIR, 1 ATM, S = 0.3
(A) PERCENT CARBON-TOTAL SAMPLE ,TA|R=533OK
1.0
O
cr
g
j-
to
O
a.
_j
<
O
0.5
-0.5
-1.0
(B) PERCENT CARBON-TOTAL SAMPLE ,TAIR=75QQK
1.0,
-0.5 I-
-1.0
(C) PERCENTAGE OF HYDROCARBONS VAPORIZED ,TA|R=533OK
1.0,
O
-------
FIG. 24
(A) TA|R=533K
l.Or-
0.5
o
DC
cr
O
in
O
a.
Q
<
cr
-0.5
-1.01-
TIME-AVERAGED NO DISTRIBUTIONS
N0.2 FUEL OIL/AIR, S=0.3, 1 ATM, 0=0.65
20 .50
100
100
1
0
^\
"~--20
1 1
2 4
80
"^50
1
6
1 |
8 10
I
12
1 |
14 16
AXIAL DISTANCE, X/D
(B) TA|R=750K
O
oc
oc
in
O
a.
Q
<
ac
0.5
-0.5
-1.0
6 8 10
AXIAL DISTANCE, X/D
12
14
16
42
-------
Spray Characteristics
As noted previously, the characteristics of the fuel spray play a determining
role in establishing the emission characteristics of a spray combustion device.
Prom the standpoint of achieving low emissions of nitrogen oxides and carbon
monoxide it is desireable to obtain a homogeneous mixture of gaseous fuel and air
where the local stoichiometry is everywhere lean such that the resulting flame
temperature is low enough to reduce the NO formation rate but high enough to pro-
mote complete oxidation of CO. This implies that the injection device must distri-
bute the fuel throughout an appropriate volume of air, and must finely atomize the
fuel to obtain high surface to volume ratios such that the liquid fuel will quickly
vaporize and diffuse through the surrounding oxidizing atmosphere. Attainment of
a high degree of atomization is crucial to the attainment of rapid vaporization; in
addition, however, high vaporization rates also may be obtained by the use of high
volatility fuels, high air temperatures, high turbulence levels, and by flow
situations where large differences exist between the mean fuel velocity and the
mean air velocity. The combustor pressure level affects vaporization rate by
changing the velocity and by raising the temperature at which droplet boiling
occurs. In the current program many of the aforementioned parameters were varied
and the influence on the spray pattern, the state of vaporization, and the
production of nitric oxide was observed. In the following paragraphs these obser-
vations are reported and comparisons between observed behavior and anticipated be-
havior are drawn.
The pressure-atomizing swirl type injector (Fig. A-l) used in this program
imparts a tangential velocity component to the liquid with the result that a
hollow-cone spray is formed. When such an injector is employed, it is important
that the cone angle be sufficiently large that the fuel is spread throughout the
surrounding air but not so large that the fuel impinges on the combustor walls.
Measurements of the spray cone angle for various operating conditions in these
tests were obtained from both the laser velocimeter data and the holographic data.
In the case of the laser velocimetry system, spray surface coordinates were taken
to be those coordinates where the data rate was highest in the case of unseeded
flows (see Appendix C) . The spray cone angle was determined by plotting these
coordinates as measured in a horizontal plane as a function of axial location and
measuring the angle between lines faired through the data points (Fig. 25). In
the case of the holographic system, the spray angle was determined from the best
fit between the surface coordinates obtained from a reconstructed image of the
spray and surface coordinates of a right circular cone (Appendix K). Tabulated
data giving the spray cone angle under both burning and nonburning conditions is
given in Table 3. The estimated error in the derived cone angle as obtained from
these measurements was approximately two to three degrees for most of the cases
examined. Excellent agreement between measurements made with both systems^was
obtained. Examination of the data indicates that the spray cone angle varied
between 64 and 69 degrees for both iso-octane and No. 2 distillate oil at the
43
-------
FIG. 25
MEAN SPRAY TRAJECTORIES
OBTAINED BY LASER VELOCIMETRY
O ISO-OCTANE, S = 0.3, 1 ATM
E ISO-OCTANE, S = 0.6 , 1 ATM
A NO. 2 FUEL OIL, S = 0.3 , 1 ATM
-1.0
-0.5
O
-------
TABLE 3. SUMMARY OF SPEAY DATA OBTAINED USING- HOLOGRAPHIC SYSTEM
-P-
Ul
Run
129-8
129-7
130-3
130-U
130-7
130-8
133-1
133-3
132-2
132-U
131-^
131-5
Combustion
No
Yes
No
Yes
No
Yes
No
Yes
No
Yes
No
Yes
Fuel
Iso-octane
Iso-octane
No. 2 Oil
No. 2 Oil
No. 2 Oil
No. 2 Oil
Iso-Octane
Iso-Octane
Iso-Octane
Iso-Octane
Propane
Propane
Pressure
(atm)
1.0
1.0
1.0
1.0
1.0
1.0
1.0
1.0
3.1
3.1
1.0
1.0
Temperature
576
57^
739
758
523
532
5U1
537
551
528
532
532
Swirl
No.
0.3
0.3
0.3
0.3
0.3
0.3
0.6
0.6
0.3
0.3
0.3
0.3
Cone Angle
degrees
81
69
68
57
75
65
71
6h
N/A
67
-
-
Mean Drop -
dia. (y)
118
105
1UU
138
N/A
140
128
120
N/A
105
-
_
-------
low inlet temperature condition. A change in the air swirl had no discernible
effect on the spray trajectory in the region where these measurements were made
(0 < X/D < 0.6). A change in pressure level within the combustor from one to
three atmospheres did not appear to have a significant effect on spray distribution.
A noticeable change in spray angle occurred only in the case when the inlet air
temperature was raised from 520K to 760K or when making a transition from a
noncombusting flow condition to a combusting flow condition. No firmly documented
explanation for the influence of inlet temperature is available. A possible
explanation for this observation is that in higher temperature or combusting flows
the dynamic pressure of the gases (pV ) is greater and therefore the increased
drag forces tend to produce a somewhat flatter droplet trajectory, resulting in a
narrower cone angle. The most dramatic change in the character of the spray
occurred when liquid propane was used as the injectant. In this case the spray
emerged from the injector as a columnar jet; no droplets were discernible within
this jet. This behavior is probably associated with the flashing of the propane
fuel as the fuel pressure decreases from the injector pressure to the combustor
pressure.
The value of the mean droplets size (Sauter mean diameter) within the spray
was determined by measuring the diameters of individual particles in the image
of the reconstructed holograms. The mean droplet diameters were found to range
from 100 to 150 microns. The results indicate that only the fuel type had a
significant influence on droplet diameter; mean droplet sizes evaluated for the iso-
octane sprays ranged from 105 to 128 microns, whereas the size range for No. 2
fuel oil sprays was from 138 to 144 microns (Table 3). Mean droplet size data were
obtained both under combusting and noncombusting flow conditions; the droplet size
in the combusting flow cases was found to be slightly smaller than for the non-
combusting flow cases but the measured size difference is too small to be considered
significant. Qualitative information on the distribution of droplet sizes within
the spray cone are available from the laser velocimetry data. Because the strength
of the signal received from the spray increases monotonically with the size of the
scattering particles (particle visibility) it is possible to obtain size informa-
tion by signal strength threshold level below which returned signals are rejected
by the data acquisition system. Size information obtained in this manner is given
in Fig. 26. As can be seen in this figure, at high threshold levels, where
only the velocity of the large droplets are recorded, two major peaks (which
correspond to the time-average spray location) are recorded as the combustor diam-
eter is traversed. When the threshold level is reduced such that signals from
the smaller droplets are accepted, the maximum droplet number densities are sig-
nificantly higher and double peaks appear which suggest that there are large num-
bers (clouds) of smaller droplets surrounding the main trajectory. The existence
of these smaller droplets adjacent to the main spray trajectory is consistent with
the fact that the smaller droplets can be affected by the turbulent gas velocity
fluctuations and therefore will tend to spread more rapidly into the surrounding
gas than will the larger droplets.
46
-------
EFFECT OF SIGNAL THRESHOLD LEVEL ON DROPLET NUMBER
FIG. 26
CO
Q
CO
r>
o
8
80
60
40
20
-1.0
X/D = 0.25
D HIGH THRESHOLD
O INTERMEDIATE THRESHOLD
DENSITY PROFILE
1.0
RADIAL LOCATION, R/R0
47
77-05-310-1
-------
With respect to the vaporization of the sprays, it is anticipated that the
fuel oil sprays will persist for a longer period of time or, equivalently, through-
out a greater spatial extent of the combustor than will the iso-octane sprays
because of the lower volatility of fuel oil. This expectation is confirmed by
both the laser velocimetry data and the phase-discriminating probe data. Com-
parison of the 10 mps contours constructed from the LV data for the fuel oil
(Fig. 27b) and the iso-octane (Fig. 28b) sprays shows the greater spatial extent
of the fuel oil spray. Because the gas phase velocity profiles for these two tests
(Figs. 27a and 28a) are quite similar and because the initial liquid velocities
were nearly the same, it can be argued that it is the decreased rate of vaporiza-
tion of fuel oil which is responsible for the greater contour dimension. With
respect to the phase discriminating probe data, comparison of the data presented
in Fig. 29 in the region to which the spray cone penetrates (R/RO = 1.0, -1.0;
X/D = 1.0) indicates greater concentrations of iso-octane were found than No. 2
fuel oil, but that the iso-octane exists in the gaseous form whereas significant
amounts of liquid fuel oil persists (Fig. 30).
The fact that detectable amounts of liquid hydrocarbons were measured in the
vicinity of the combustor wall raises the question as to whether the spray persists
for a sufficient period of time that appreciable quantities of liquid fuel pene-
trate to the wall. The magnitudes of the percent carbon in the total samples as
determined from the phase-discriminating probe indicate that this was not the case.
Noting that the level of unburned hydrocarbon in the liquid phase can be determined
by multiplying the total percentage of unburned hydrocarbons (e.g., Fig. 29b) ,
by the percentage unvaporized (Fig. 30b), it is determined that the largest value
of the percent carbon (0.5) due to the existence of liquid fuel is found down-
stream of the anticipated point of impingement as predicted from spray trajectory
data and is small in magnitude compared to that value corresponding to stoichio-
metric mixtures (13.2 for iso-octane). Furthermore, no buildup of carbonaceous
material on the cooled combustor wall was detected nor was any evidence of streak-
ing or staining of the combustor walls noted during the periodic inspections of
the test apparatus.
As noted above, the results of the phase-discriminating probe tests indicate
that liquid fuel exists in the downstream region of the combustor; that is, down-
stream of the region associated with the spray cone itself. Existence of liquid
fuel in this region is expected since the largest droplets require long burning
times and will thus travel significant distances before being consumed. Again,
one would expect to find more liquid fuel in the case of the tests conducted with
fuel oil than in the case of tests with the more volatile iso-octane. Comparisons
of the unburned hydrocarbon data for the two different fuels (Fig. 29) indicate
that no major differences in the spatial pattern of the total (liquid plus vapor)
unburned hydrocarbons exists and, surprisingly, somewhat higher levels of unburned
hydrocarbons exist for iso-octane than for fuel oil. This may well be due to the
differences in reaction rates associated with differences in liquid fuel penetra-
tion into the air stream. According to expectations, the fraction of the unburned
48
-------
MEAN AXIAL VELOCITY DISTRIBUTIONS
NO. 2 FUEL OIL/AIR, SW,RL = 0.3, 1 ATM, TA|R . 53^, 0= 0.65
FIG. 27
(A) GAS PHASE
-1.0 r
o
cr
cr
z~
O
o
O
<
Q
cc
-0.5 -
0-25 0.50
AXIAL DISTANCE, X/D
(B) DROPLET
-1.0
O
EC
0.75
0.25 0.50
AXIAL DISTANCE, X/D
0.75
49
-------
MEAN AND RMSAXIAL VELOCITY DISTRIBUTIONS
ISO-OCTANE/AIR, 5 = 0.3,1 ATM, TA,R = 533°K, 0 = 0.65
FIG. 28
(A) GAS PHASE
-1.0
URMS>10M/SEC
(B) DROPLET
-1.0 r-
O
en
-z.
u
o
<
D
-0.5
0.5
1.0
RECIRCULATION ZONE
U<10 M/SET
SPRAY TRAJECTORY
(C) DROPLET
-1.0
O
cc
o
CJ
o
<
Q
-0.5 h
0.25 0.50
AXIAL DISTANCE, X/D
0.75
50
-------
FIG. 29
TIME-AVERAGED DISTRIBUTIONS OF TOTAL UNBURNED HYDROCARBONS
S = 0.3, 1 ATM, TA|R = 533°K
(A) ISO-OCTANE
1.0
-1.0
(B) N0.2 FUEL OIL
o
oc
DC
z"
o
CO
O
O.
_)
<
D
<
cc
0.5
-0.5
-1.0
0 1
34567
AXIAL DISTANCE, X/D
51
79_01-78-3
-------
TIME-AVERAGED PERCENTAGE OF HYDROCARBONS VAPORIZED
S = 0.3, 1 ATM,TA|R = 533° K
(A) ISO-OCTANE
FIG. 30
O
cc
C/3
O
Q.
<
Q
<
CC
1.0
0.5
-0.5
-1.0
I I I I
J I I
(B) N0.2 FUEL OIL
1.0
O
DC
£ 0-5
O
Q.
Q
< -0.5
cc
-1.0
012345678
AXIAL DISTANCE, X/D
52
79-01-78-4
-------
fuel that is in the liquid state is indeed higher for the fuel oil case than
in the iso-octane case (Fig. 30).
Data obtained for propane, Fig. 31, show extremely high levels of unburned
hydrocarbons near the combustor axis which is a result of aforementioned poor
fuel distribution obtained with propane. The existence of significant amounts of
liquid propane in the downstream region is indicative of the fact that the high
fuel concentrations cause a reduction in the local gas temperature which precludes
further vaporization until sufficient aerodynamic entrapment of the hot surround-
ing flow occurs.
The effect of a change in inlet temperature from 530K to 750K on the distri-
bution of hydrocarbons is illustrated by the contour patterns in Figs. 29b, 30b,
and 32. Again, it can be seen that .no major change in the pattern of unburned
hydrocarbons occurred; however, as expected, the fraction vaporized was somewhat
higher in the higher temperature case.
Another feature of the fuel spray which bears comment is the location of the
spray relative to the locations of the low velocity regions which act as flame
stabilization sites. In highly stable combustion systems employing pressure
atomizing fuel nozzles, the fuel is injected directly into large regions of recir-
culatory gas flow located on the combustor axis downstream of the fuel injector
(Ref. 13). Typically, the bulk of the injected spray passes through the recircu-
lating flow while a fraction of the smaller droplets remain in the recirculating
flow and combine with entrained oxidizer thereby maintaining a hot gas core which
serves as a flame stabilization region. As was discussed previously, particle
visibility problems precluded velocity measurements in the vicinity of the spray
at low swirl and, consequently, no time-mean recirculating flow is shown. However,
a torroidal-shaped recirculating flow was detected in the high swirl case and
positive gas velocities existed at all points on the combustor axis (see Figs. 28
and 33). Curiously, local regions of negative mean droplet velocity existed in
the low swirl case (as well as in the high swirl case) even though the time-mean
gas velocity was everywhere positive (see Figs. 27b and 28b) . It is also noted
that very large fluctuations in the velocity of the droplets occur within this
device under both the moderate and low swirl conditions (Figs. 28c and 33c). These
observations together with the previously reported transient pressure and luminosity
fluctuations confirm that transient aerodynamic phenomena are influencing the
behavior of the flow in the vicinity of the injector and that the application of
the frequency tracking laser velocimetry techniques will be required to gain a
thorough understanding of the detailed time-dependent flow processes.
The laser velocimetry system was also used to measure the tangential com-
ponents of the gas velocity and of the fuel velocity (see Fig. 34). The fact that
the sense of rotation of the air (solid symbols) is the same as that of the fuel
(open symbols) is evident in the figure; also evident is the increased tangential
velocity component of the gas in the moderate swirl case. Note that in the
53
-------
Fie:. 31
TIME-AVERAGED DISTRIBUTIONS OF UNBURNED HYDROCARBONS
PROPANE,! ATM, TA)R=533°K, S= 0.3
(A) PERCENT CARBON-TOTAL SAMPLE
1.0 i
O
cc
tr
g
H
CO
O
Q_
_l
<
Q
0.5
-0.5
-1.0
(B) PERCENTAGE OF HYDROCARBONS VAPORIZED
O
a:
Z
O
H
CO
O
Q.
_1
<
Q
<
tr
1.0
0.5
-0.5
-1.0,
AXIAL DISTANCE, X/D
54
-------
TIME-AVERAGED DISTRIBUTIONS OF UNBURNED HYDROCARBONS
' 32
NO. 2 FUEL OIL, 1 ATM, TA|R = 750° K, S = 0.3
(A) PERCENT CARBON-TOTAL SAMPLE
1.01
O
-------
FIG. 33
MEAN AND RMS AXIAL VELOCITY DISTRIBUTIONS
ISO -OCTANE/AIR. S = 0 6, 1 ATM. TA |R 533°*.* 065
4ii RECIRCULATION ZONE
(A) GAS PHASE
-1.0,
SSfe: UR|u|c>10 M/SEC
<20M/S u = 20 M/SEC
O
-------
FIG.
MEAN TANGENTIAL VELOCITY PROFILES
X/D = 0.123
T 60 VELOCITY (M/SEC)
O ISO-OCTANE, S = 0.3
D ISO-OCTANE, S = 0.6
A NO. 2 FUEL OIL, S = 0.3
SOLID SYMBOLS DENOTE AIR SEED 4" 40
4-20
-1.0
RADIAL POSITION,
57
76-10-29-9
-------
moderate swirl case vaporization of fuel spray should be enhanced by the large
velocity differences which exist between the liquid and gas flow. Measurements
taken at a second downstream station (X/D = 0.335 vs X/D = 0.123) indicate that
a slight decay of the velocity difference between the gas and fuel droplets has
occurred; however, the fuel and droplets still maintain distinctly different tan-
gential velocities (Fig. 35).
To summarize these observations, the holographic and laser velocimeter measure-
ments indicate, that within the region close to the injector (within one combustor
diameter downstream of the injector tip) the fuel spray was observed to maintain
the structure of a hollow cone and to penetrate throughout the cross-section
of the swirling airflow. Large differences between air and fuel time-mean veloc-
ities exist in this region as do large fluctuations in the local gas and droplet
instantaneous velocities. These velocity field characteristics probably influence
the vaporization and consumption of the smaller droplets in this region and thereby
affect flame stability, but the impact of the velocity fluctuations on the larger
droplets constituting the major portion of the fuel spray is undetermined. The
phase-discriminating probe data indicate that significant amounts of liquid fuel
exist for at least the first five combustor diameters, particularly near the outer
radii of the combustor. The initial mean droplet size of the fuel oil spray is
larger than that of the iso-octane spray; this together with the lower vapor pres-
sure of the fuel oil would result in longer droplet burning times and hence are
probably responsible for the greater concentrations of nitric oxides produced in
some of the tests conducted with fuel oil.
58
-------
MEAN TANGENTIAL VELOCITY PROFILES
X/D = 0.335
^ ISO-OCTANE, S = 0.3
Q ISO-OCTANE, S = 0.6
A NO.2 FUEL OIL, S = 0.3
SOLID SYMBOLS DENOTE AIR SEED
FIG. 35
VELOCITY (M/SEC) _. 40
RADIAL POSITION
59
76-10-29-2
-------
SECTION IV
RECOMMENDATIONS
The experimental investigations carried out under EPA contracts 68-02-1092
and 68-02-1873 have shown that variation in inlet conditions, e.g., pressure,
inlet air swirl and inlet air temperature, produce major changes in the mean flow
field, including vaporization and mixing rates, within a liquid fuel turbulent
diffusion flame burner which result in subsequent changes in energy release rates
and pollutant formation and destruction. The variation in pollutant emission
trends with fuel type, found in the present investigation, serves to illustrate
the difficulty in obtaining general relationships between pollutant emissions and
inlet conditions in liquid fuel combustors. Pollutant formation and destruction
are intimately connected with the flow field structure within the combustor and
the structure depends on the spray pattern and on the interaction of the spray
with the gas flow. The present data base is inadequate to permit definitive
correlations of the flow field structure and pollutant emissions with burner
inlet conditions.
Additional data on the effects of changes in inlet conditions on the mean
and fluctuating flow field structure of liquid fuel turbulent diffusion flames
and the subsequent effects on pollutant formation and destruction are required.
Particular emphasis should be placed on determining spray characteristics,
including droplet trajectories and size distributions, and in measuring fuel
vaporization rates. Existing optical techniques for spray visualization in com-
busting flows should be refined to permit resolution of fuel droplets in the 5-50
Um range for spray particle densities of interest in real combustion devices to
assist in these measurements. For each liquid fuel examined, separate fundamental
experiments should be used to shed light on (a) the thermal decomposition of the
fuel and (b) the combustion of droplet arrays. The complementary droplet combus-
tion experiments would be particularly useful in examining the combustion character-
istics of fuels containing bound nitrogen.
Recent studies of pollutant emissions from liquid fuel combustors have
shown that changes in operating conditions which produce decreases in NO emissions
generally result in significant increases in particulate emissions. Results from
the present investigation followed similar trends. These observations suggest
that future studies should be concerned with the effects of variation in operating
conditions on formation and destruction of multiple pollutant species, including
particulates (size, number density and composition) and various amine and cyano
compounds.
60
-------
APPENDIX A
DETAILS OF EXPERIMENTAL APPARATUS AND INSTRUMENTATION
Combustor Facility
A conventional pressure-atomizing swirl-type nozzle, Fig. A-la, which
produced a nominal 60 deg hollow-cone spray with a nominal droplet Sauter mean
diameter of lOOym in quiescent air at atmospheric pressure was used for iso-
octane and No. 2 distillate fuel. A tangential-feed pressure-atomizing nozzle,
Fig. A-lb, with a similar spray angle and in which the full pressure drop occurred
across the exit orifice was used with propane to maintain the fuel liquid to the
point of injection. The direction of rotation imparted by the swirlers to both
the fuel and the air streams were identical for each of the configurations tested.
The air swirl vane designs, shown together with the fuel injector assembly
in Fig. A-2, are similar to those used previously in the gaseous fuel test pro-
gram (Ref . 10). The swirl number, S, was computed from the injector geometry Z,
and the angle of the swirl vanes, ri, according to the following expression
(Ref. 11):
(A-l)
The swirl number is simply the ratio of the angular momentum flux to the axial
momentum flux multiplied by an effective nozzle diameter.
The 12.23-cm diameter, 100-cm long instrumented combustor is divided into
five water-cooled zones of approximately equal length. Water flow can be set
independently in each zone, as needed, to keep wall temperature (^500°K) roughly
constant along the entire length of the combustor. Wall temperatures are set
and monitored using thermocouples installed on the outer surface and at various
depths in the combustor wall and cooling passages. Static pressure taps are
also installed at several locations along the combustor. Flow exhausts from
the combustor and extender sections to the facility exhaust stack. Combustor
extender pieces, 33.4 cm in length, are inserted when required to fully contain
the flame; the extender section consisted of two extender pieces during all of
the current experimental effort.
Gas Sampling and Temperature Probes
Species concentration distributions within the combustor were measured
using a traversing gas sampling probe, an exhaust gas sampling rake and a
61
-------
PRESSURE ATOMIZING LIQUID FUEL INJECTORS
FIGA-1
FILTER
SWIRL BLOCK-7
ORIFICE
(a) AXIAL SWIRL
TANGENTIAL
FEED
ORIFICE
CM
(b) RADIAL SWIRL
62
76-09-41-1
-------
INJECTOR AND SWIRL VANE GEOMETRIES
FLOW'
I I I I I I I I I I I I I I I I I II I I I I I
II
i /77777/77/7/7///7/r/ / / / /
VANE 0.163 CM
THICK 316 SS
1.43 CM INJECTOR STEP
3J
g
Z = dh/d
0.203
L(CM) S
3.732 O.3
0.6
T?(DEG)
40
60
NO. OF
VANES
12
8
X(CM)
2.24
-------
phase-discriminating sampling probe. Composition information is determined
on-line by aspirating flow through the cooled probes and analyzing the gas
sample using a Scott Model 119 Exhaust Gas Analyzer and a heated discrete-
sampling hydrocarbon analyzer. Pressurized hot water at 400°K was used as
the probe coolant to minimize wall-catalyzed reactions and to prevent water
and fuel condensation and loss of species within the sampling lines.
The exhaust probe rake, located at the exit of the extender section, consists
of five identical probes centered on equal area annuli (Pig. A-3). The individual
probes are manifolded downstream and a single mixed sample is transferred to
the gas analyzer. Radial traverses are made at selected axial locations within
the instrumented combustor section using a single gas sampling probe of similar
design (Fig. A-4). The inlet flow into both sampling probes was maintained choked,
resulting in aerodynamic cooling of the sample by means of a rapid internal ex-
pansion. This expansion combined with the wall cooling effect served to quench
chemical reactions involving stable species.
Temperature profiles at the exhaust plane and within the combustor were
measured by traversing a calibrated-heat-loss thermocouple probe across a
combustor diameter. Although conventional thermocouple materials limit applica-
tion of these sensors to temperatures below about 2000°K, cooling the exposed
junction by conduction heat transfer extends the range of thermocouple utiliza-
tion above the melting point of the material to the 2000-2500°K range. In order
to obtain the local stream temperature, the measured stream thermocouple tempera-
ture must be corrected for conduction and radiation heat losses; therefore, cali-
bration information is acquired simultaneously with the required temperature
measurement. The probe consists of three thermocouples including an iridium -
10 percent rhodium/iridium thermocouple which protrudes from a water-cooled
copper base into the reacting flow, and two platinum - 10 percent rhodium/platinum
thermocouples installed on the ends of the iridium wire to record the base tempera-
ture and thereby permit calculation of the conduction heat loss (Fig. A-5). A
thermocouple probe of this type was applied without difficulty in the combustion
environment of the present program. Confidence in the accuracy of the temperature
measurements was established in the previous contract effort by measurements
made at identical test conditions using a conventional thermocouple probe and a
double-sonic-orifice probe (Ref. 10).
Gas Sampling System
The gas samples withdrawn through the five-probe exhaust rake or the traversing
probe are analyzed on-line to determine the time-averaged concentrations of carbon
dioxide (C02), carbon monoxide (CO), oxygen (02), nitrogen oxides (NO, N02) and
unburned hydrocarbons (THC). The samples are transferred to the analytical
instruments through a teflon-coated, flexible line which is heated (^ 400°K)
64
-------
EXHAUST SAMPLING PROBE RAKE
PROBES CENTERED ON EQUAL AREAS
FLOW FROM ALL PROBES MIXED BEFORE
ON-LINE ANALYSIS
ON
FLOW
ALL DIMENSIONS IN CM
NOT TO SCALE
15.88
I3
HIGH PRESSURED I _
FLOW
_ __ .
0.198 0.08o L_ 0.025
!-
-------
TRAVERSING GAS SAMPLING PROBE
1.27 D-
0.20 D
0.95 D
0.31 D
^
TIP DETAILS
45.72
g
-j
ID
I
ALL DIMENSIONS IN CM
2.54
-------
CALIBRATED-HEAT-LOSS THERMOCOUPLE PROBE
-FLOW
WATER OUT
o
i
WATER IN
TRAVERSING MECHANISM
I I I I I I I I I I L
Es j 41..
CM l 2
P
I
en
-------
electrically to prevent water condensation. The sample is then directed through
a condensate trap (^ 277°K), where most of the water is removed, and it is pumped
through an unheated, teflon coated, aluminum line to a Scott Model 119 Exhaust
Analyzer. A schematic diagram of the sampling system is shown in Fig. A-6-
A stainless steel bellows pump increased the sample pressure from subatmospheric
levels to 1 atm as required by the Exhaust Analyzer. The Analyzer, located in
the combustion facility control room, approximately 10 m from the combustor, was
used to measure the molar concentrations of CO, CO., (^, NO, N02 and THC.
The Scott Model 119 Exhaust Analyzer, (Fig. A-7), is an integrated system,
with flow controls for sample, zero and calibration gases conveniently located
on the control panel. The incoming gas sample passes through a refrigeration
condenser (^ 275°K), to remove residual water vapor. As the sample passes from
the condenser, it is filtered to remove particulate matter. The Exhaust Analyzer
is comprised of five different pieces of analytical instrumentation. Beckman
Model 315B Nondispersive Infrared (NDIR) Analyzers were used to measure the CO
and C02 concentrations (mole fractions) in the gas sample. Concentration ranges
available on the CO analyzer were from 0-200 ppm to 0-15 percent on several
scales. Concentration ranges available on the C02 analyzer were 0-4 percent
and 0-16 percent. The accuracy of the NDIR analyzers is nominally +_ 1 percent
of full scale. A Scott Model 125 Chemiluminescence Analyzer was used to measure
the NO and N02 concentrations in the gas sample. Concentration ranges available
with this instrument were from 0-1 ppm to 0-10,000 ppm on several scales, with
a nominal +_ 1 percent of full scale accuracy. The thermal converter used in the
chemiluminescent analyzer was stainless steel, and was operated at a temperature
of approximately 1030°K. The converter efficiency (i.e., percent NO- dissociated)
was determined using the method outlined in Refs. 10 and 15. In the present study
a converter efficiency of 99 percent was measured, with an uncertainty in the
measurement of 4 percent. A Scott Model 150 Paramagnetic Analyzer was used to
measure the Q£ concentration in the gas sample. Concentration ranges available
with this instrument were from 0-1 percent to 0-25 percent on several scales,
with a nominal accuracy of + 1 percent of full scale. A Scott Model 116 Total
Hydrocarbon Analyzer was used to measure the hydrocarbon concentration in the
gas sample. This analyzer utilizes an unheated flame ionization detection
system to provide for measurement of hydrocarbons (as carbon) in concentration
ranges from 0-1 ppm to 0-10 percent, with a nominal accuracy of +_ 1 percent of
full scale. The unheated analyzer was used to measure exhaust hydrocarbons only
for propane. Output signals from the various analyzers are displayed on chart
recorders. The Analyzer was calibrated prior to each test by flowing zero
gases and calibration gas mixtures having compositions known to within one per-
cent. Typically, at each test point, sampling data were acquired for a period
of 2-4 min.
68
-------
FIG.A-6
SCHEMATIC DIAGRAM OF ON-LINE GAS ANALYSIS SYSTEM
ASPIRATED GAS SAMPLE
CALIBRATION AND
ZERO GASES
1
(
(
<
(
CO
TF
r r - r )
NDENSA
1AP (277°
\ir i i 1-1
VLS/Z2.
ETERE
rs
K)
r r.r r v/
HEATED LINE
m
BELLOWS
PUMP
*=^J==
II
SCOTT MODEL 119
EXHAUST
ANALYZER
(CO,CO2,NOX,02,THC)
69
N12-161-1
-------
EXHAUST GAS ANALYTICAL SYSTEM
co NDIR rr
ANALYZER
-J
01
CO2 NDIR
ANALYZER"
NOX
CHEMILUMINESCENT
ANALYZER
FID
HYDROCARBON
ANALYZER
PARAMAGNETIC
ANALYZER
p
>
i
-------
Phase-Discriminating Sampling Probe
The mole fractions of the vapor fuel and of the total (liquid + vapor) fuel
within the combustor were measured by traversing a cooled phase-discriminating
probe across the combustor diameter at selected axial locations downstream of
the fuel nozzle. The probe designs employed were based on the work of Wadleigh
and Oman (Ref. 16) who investigated the effects of the probe geometric design
parameters and flow rates on the probe performance. Previously, uncooled phase-
discriminating sampling probes have been designed, built and employed to measure
the state of vaporization of nonburning fuel sprays (Ref. 17). The current study
represents the first attempt to employ this probe design in a combusting flow
field.
A schematic diagram showing the construction of the tip of the cooled phase-
discriminating probe employed in this study is given in Fig. A-8. The central
passage acts as a conventional gas sampling probe and collects the total (liquid +
vapor) sample. Isokinetic flow is established within this passage by adjusting
the flow rate through the tube so that the static pressure close to the tube lip
is equal to the combustor static pressure. A tube oriented perpendicular to the
axis of the total-sample tube is used to extract the vapor sample. A suction
tube surrounding the vapor tube is used to purge any liquid which collects on
the surface of the total-sample tube and which otherwise would spill over into
the vapor sample tube and contaminate the vapor sample. Water jackets surround
the sampling tubes so that the collected samples can be quickly quenched. Also,
provisions were incorporated into the design for introducing a flow of nitrogen
into the sample close to the probe tip in order to quench vaporization and chemical
reaction. This purge feature was not employed in the present study. Water
cooling also is required to ensure structural integrity of the probe. Heat trans-
fer analyses of the probe design indicate that the probe tip can withstand gas
temperatures of 2500°K and pressures of 7 atmospheres for a water flow rate of
0.3 kilograms per second. A water supply pressure of 4 x 105 Newt on s/meter2 was
adequate for providing this flow rate. It further is required that over-cooling
of the probe be avoided since this would lead to condensation of the fuel vapor
samples. Temperatures of the samples were monitored by a thermocouple inserted
into the total-sample line at a location 15.2 cm from the probe tip (Fig. A-9) .
The probe coolant flow rate was regulated so that the measured temperature was
maintained above 560°K.
A photograph showing the phase-discriminating probe assembly is given in
Fig. A-10. The probe is held in position by a water-cooled mounting plate
which can be installed in any of the window ports in the combustor test se
The probe is traversed by a hydraulic actuating mechanism which is equipped
a slide wire probe position indicator.
71
-------
FIG.A-8
SCHEMATIC DIAGRAM OF PHASE-DISCRIMINATING PROBE TIP
TOTAL SAMPLE
PROBE COOLANT WATER
ISOKINETIC SAMPLING
PRESSURE TAP ,. SUCTION FLOW
VAPOR SAMPLE
BURNING SPRAY
ALL DIMENSIONS IN CM
V SUCTION - 0.236 O.D. x 0.020W
\
TOTAL - 0.556 O.D. x 0.041W
\
\ r- PRESSURE-0.159 O.D.
x 0.023W
-VAPOR -0.1 57 O.D. x 0.025W
\
1.905
O.D.
72
76-06-145-3
-------
SCHEMATIC DIAGRAM OF PHASE DISCRIMINATING PROBE
ALL DIMENSIONS IN CM
FIG.A-9
NITROGEN QUENCH SYSTEM
SUCTION TUBE
0.019
VAPOR-SUCTION
TUBE SECTION
»- COOLANT WATER
OUTLET
COOLANT WATER
INLET
TOTAL SAMPLE
73
76-07-202-1
-------
FIG. A-10
PHASE-DISCRIMINATING PROBE ASSEMBLY
INLET
MOUNTING PLATE
COOLANT SYSTEM
OUTLET
INLET
PROBE COOLANT
SYSTEM
HYDRAULIC
ACTUATOR
HEATED
SAMPLE LINJS
OUTLET
RL76-189-A
SAMPLING PROBE
INLET
WATER COOLED
MOUNTING PLATE
- SLIDE WIRE
PROBE POSITION
INDICATOR
THERMOCOUPLE
SAMPLE TEMPERATURE
76-06-169-1
-------
Fuel Analysis System
The concentration of fuel in the samples withdrawn by the phase-discriminatin*
probe is obtained from on-line determinations of the hydrocarbon content using a
flame ionization detector. A schematic diagram of the sampling system is shown
in Fig. A-ll. The sample is pumped from the probe through electrically-heated
stainless steel lines, which maintain a sample temperature of approximately 500°K,
and through glass wool particulate filters to the gas analysis equipment. Sample'
flow rates are controlled using electrically-operated throttling valves. Flow
rates are determined from orifice pressure measurements. The pressures upstream
and downstream of the orifices are measured by use of a Validyne DP7 Diaphragm
Magnetic Reluctance Transducer employing a +3.4 atm diaphragm. The transducer used
in establishing the pressure differential between the total sample tube and the
combustor is a similar unit having a + 7 x 10 atm diaphragm. The flows then
pass through MB-158 High Temperature Welded Bellows Pumps which increase the
sample pressure from subatmospheric levels to one atm as required by the analyzer
system. Stainless steel bellows pumps are used to prevent contamination of the
samples. The metal bellows pumps are contained in ovens equipped with temperature
limit switches which only permit operation of the pumps at the elevated tempera-
tures required to prevent sample condensation. Because the flow fate in the
total sample line required to ensure isokinetic conditions at the probe tip is
greater than the flow capacity of the metal bellows pump, a bypass pump is
connected to the total sample line upstream of the bellows pump. The flow rate
through this CAST Model 0522 rotary vane vacuum pump is controlled by an
electrically-operated throttling valve. Flow is passed around the metal bellows
pumps when operating the combustor at above-atmospheric conditions. The vapor
and total samples are passed to the ionization gauge via a ten-port sampling
valve. Pressure relief valves having a 2 x 10"^ atm cracking pressure are in-
stalled in the lines upstream of the sampling valve in order to ensure atmospheric
pressure in the ionization gauge. The flow in the suction line, which is not
analyzed, is transferred to an exhaust line.
The ten-port sampling valve permits uninterrupted flow of a carrier gas, argon
and gas samples through the ionization gauge (Fig. A-12) . Conventional hydro-
carbon analyzer systems employing flame ionization detectors provide a continuous
sampling capability, but for the purposes of analyzing fuel sprays where locally
high fuel/air ratios exist, a method of preventing saturation of the ionization
gauge must be provided. The sampling valve employed in this system is designed
to extract small discrete samples from either of the transfer lines and to dilute
those samples with carrier gas prior to delivery to the ionization gauge. As
shown in the figure, as the valve slide is repositioned a discrete amount of
either the total fuel sample or vapor fuel is injected into the carrier gas tor
delivery to the ionization gauge. The peak concentration of the sample Caching
the ionization gauge is determined by the length and size of the loops on tne
sampling valve and the length of the line between the sampling valve and
ionization gauge.
75
-------
SCHEMATIC DIAGRAM OF PHASE DISCRIMINATING PROBE GAS ANALYSIS SYSTEM
EXHAUST
PHASE DISCRIMINATING PROBE
ELECTRICALLY
HEATED SAMPLE
LINES
WATER- COOLED
PROBE LINES
P P
T
o
zn
THROTTLING
VALVES
ORIFICES
PURGE
SYSTEM
CALIBRATION
GASES
»- TO ELECTROMETER
TEN PORT
SAMPLING
VALVE
i !
BYPASS LINE
ARGON
CARRIER
RELIEF
VALVES
EXHAUST
THROTTLING^
VALVE
BYPASS
PUMP
I
g
TJ
P
-------
TEN PORT SAMPLING VALVE
TWO POSITION SLIDE
SAMPLE
VAPOR
X
CARRIER
TOTAL C
/
\
tV
\
TOTAL
r
-*- EXHAUST
TO DETECTOR
X_
VAPOR
CARRIER-
TOTAL
TWO POSITION SLIDE
VAPOR
X
f
IDE 7
P/
\
/
7
-*- EXHAUST
TO
DETECTOR
EXHAUST
SAMPLE
PORT (T) CONTAINS TOTAL SAMPLE TO BE INJECTED DURING
NEXT SWITCHING OF SLIDES
PORT (2) CONTAINS VAPOR SAMPLE TO BE INJECTED DURING
NEXT SWITCHING OF SLIDES
01
o
a) SLIDE POSITION AFTER INJECTION OF VAPOR SAMPLE
b) SLIDE POSITION AFTER INJECTION OF TOTAL SAMPLE
P
j^
10
-------
A Gomac ionization gauge and signal conditioner (electrometer) are used to
measure the hydrocarbon concentration in the gas samples. The ionization gauge
is piloted with a hydrogen air flame using a 60/40 hydrogen/nitrogen fuel supply.
A photograph showing the ionization gauge and ten-port sampling valve in a con-
stant-temperature oven is given in Fig. A-13.
All of the components of the gas analysis system are installed on an instru-
mentation cart which was designed to withstand the harsh environment encountered
in combustion facility test cells. A photograph showing installation of the
equipment on the instrumentation cart is given in Fig. A-14. The instrumentation
cart is located approximately 2 meters from the base of the probe. The probe
hydraulic drive mechanism, flow control valves, and sampling valves are operated
from a control room where the output of the ionization gauge is recorded on a
strip chart recorder (Fig. A-15). The recorder displays the hydrocarbon concen-
tration as a function of time and also mechanically integrates the concentration
with respect to time. A typical trace is displayed in Fig. A-16. The system is
capable of analyzing a gas sample every 30 seconds.
Laser Holographic System
A laser holographic system was developed and used to measure the spatial
location of the fuel droplets and mean droplet size in liquid fuel sprays
emanating from the fuel injector. A schematic diagram of the off-axis holo-
graphic system used to make these measurements is shown in Fig. A-17. The beam
from a Q-switched ruby laser (Korad Model K-1QP) is separated into an object beam
and a reference beam by a beam splitter. A Tropel Model 280 expander/collimator
is used to expand the reference beam, which initially is less than five milli-
meters in diameter, to an approximately 12.5 cm diameter beam which strikes the
holographic plate. The collimator is adjusted so that the wavefronts emanating
from the lens of the collimator are plane. The reference beam is oriented so
that the angle of the reference beam with respect to the object beam is approxi-
mately 45 deg. This angle is a compromise between larger angles which would
cause the fringe patterns developed on the holographic plate to become very
closed spaced, thereby reducing resolution, and smaller angles which would
require the holographic plate to be moved away from the object. It is desirable
to keep the holographic plate close to the object in order to maintain the high
f-number optics required for good resolution and to permit the use of high f-
number .reconstruction optics which provide a short depth of focus.
The portion of the laser beam which is undeflected by the beam-splitter
passes through a lens to a diffuser (ground glass plate). The diffuser causes
the rays of light to be transmitted to the object (the spray over a large range
of angles relative to the optical axis.) The light emanating from the spray is
the true object beam, and consists of light transmitted through the transparent
78
-------
FIG. A-13
HYDROCARBON ANALYZER
VAPOR SAMPLE LINE
TEN PORT
SAMPLING
VALVE
SAMPLE LINE LOOP
CARRIER GAS LINE
TOTAL SAMPLE LINE
BELLOWS
PUMP
IONIZATION
GAUGE
R76-189-B
76-06-169-2
79
-------
FIG. A
PHASE DISCRIMINATING PROBE INSTRUMENT CART
HEATED TRANSFERS LINES
OVEN CONTAINING BELLOWS PUMP
AND'IONIZATION GAUGE
MOTOR FOR BELLOWS PUMP
, ELECTROMETER
TOTAL SAMPLEH
SUCTION SAMPLE
VAPOR SAMPLE
ORIFICE SYSTEMS
HEATING TAPE
TEMPERATURE
CONTROLLERS
76-06-185-2
80
-------
PHASE-DISCRIMINATING PROBE CONTROLS AND READ-OUT
FfG. A_1
STRIP **
CHART *?
RECORDER
SAMPLE FLOW AND
PROBE POSITION CONTROLS
REMOTE ELECTROMETER CONTROL
PRESSURE SCANNER CONTROLS
R76-189-D
76-06-169-3
81
-------
FIG.A-16
TYPICAL EXPERIMENTAL TRACE FROM HYDROCARBON ANALYZER
LU
2
I-
WINDOW3
POSITION R = 2.50
WINDOWS
POSITION R = 0.01
ATTENUATION C = 10~10
MECHANICAL INTEGRATOR-
OUTPUT
82
76-06-145-7
-------
FIG.A-17
SCHEMATIC DIAGRAM OF LASER HOLOGRAPHY SYSTEM
r-if>
DIFFUSER
MIRROR >2
FILTER
HOLOGRAPHIC PLATE
COMBUSTOR TEST SECTION
/
OBJECT .
BEAM I I
MIRROR
LENS
BEAM SPLITTER
/ /
\ \ REFERENCE BEAM
\ \
\ \
\ \
\ \
N \
\
MIRROR
EXPANDER/COLLIMATOR
He-Ne
ALIGNMENT
LASER
PRISM
MIRROR
PULSED RUBY LASER
83
76-OB-213-4
-------
droplets, reflected from the surface of the droplet and refracted by the edges
of the droplet. Light from each point in the spray reaches all points on the
holographic plate thereby providing the information required to reconstruct a
three-dimensional image of the spray. Depending on the distance traveled by
the light emanating from the spray, the object beam will either interfere with
or be reinforced by the reference beam, thereby producing fringes. The holography
system employed is of the transmission type, since light from the diffuser is
transmitted through the object to the holographic plate. With such systems
the reconstructed images of the objects appear as silhouettes if the objects
are opaque. In the case of a transparent droplet, the edge of the droplet
appears dark because light at the edge is refracted out of the field of view
of the holographic plate. Light that passes through the center of the droplet
reaches the holographic plate and, therefore, the droplet image appears as a
dark outer ring with a bright center. The dark outer edge represents the true
dimension of the transparent droplet in the object field.
To reduce fogging of the holographic plate by light emitted from the hot
combustion gases, a filter was installed between the test section window and
the holographic plate. Both a gelatin filter (Kodak Wratten No. 29) and a
high quality interference filter were employed. These filters, the transmission
characteristics of which are shown in Fig. A-18, were equally effective in
blocking the radiation emitted from the combustion gases. The holograms were
recorded on Agfa Scienta Type 10E75 film on a 4 x 5 in. glass base. A remotely-
controlled film magazine capable of holding eight film plates was utilized so
that a series of holograms could be obtained without entering the test cell or
disturbing the alignment of the optical components.
A schematic diagram showing the components used to reconstruct the holograms
is given in Fig. A-19- A Spectra-Physics Stablite Model 124A helium-neon laser
was used as the light source. The laser beam was expanded by a Tropel Model 280
expander/collimator to produce a beam of light approximately 12.5 cm in diameter.
The light illuminated the holographic plate mounted in an articulated holder to
facilitate precise orientation of the hologram relative to the incident laser
beam. Light diffracted from the hologram formed a real image of the spray. This
image was examined by eye using a 12-power loupe mounted on a tripod and rack-and-
pinion mechanism so that the loupe could be translated in three-dimensions.
The traversing mechanisms were equipped with index scales having 1-mm divisions
so that the coordinates of the object being viewed could be determined. The
loupe was equipped with a reticle having square outlines ranging in size from
100 to 250 microns. This size range was comparable with the range of sizes of
the larger droplets observed in the spray.
84
-------
FIG.A-18
FILTER TRANSMISSION CHARACTERISTICS
1.00
0.80
> 0.60
GELATIN
FILTER
(WRATTEIM NO. 29)
V)
V)
0.40
0.20
01
5600
o
6943 A
(RUBY LASER)
I
INTERFERENCE
FILTER
6000 6400 6800
WAVELENGTH (ANGSTROMS)
7200
7600
76-09-2-7
85
-------
FIG.A-19
SCHEMATIC DIAGRAM OF
HOLOGRAM RECONSTRUCTION APPARATUS
HELIUM-NEON LASER
EXPANDER/
COLLIMATOR
LOUPE AND EYE
86
.76-04-337-2
-------
Laser Velocimeter
In the gaseous fuel test program (Ref. 10), a laser velocimeter was used
to measure the mean and rms gas velocities in the initial regions of the reacting
flow. It is desirable to obtain similar data in the liquid fuel program, and
in addition it would be worthwhile to obtain data on droplet velocities and tra-
jectories. However, in liquid fuel combustors the flow field immediately down-
stream from the fuel injector is complex, and the utility of laser velocimeter
measurements in these flows is uncertain. Significant questions arise concerning
the effect of the spray on the propagation of the laser beam within the combustor
and concerning the ability of the laser velocimeter to discriminate between large
fuel droplets, which generally have velocities different from the gas stream,
and small droplets or seed particles, which move with the gas stream. Accordingly,
a series of tests were carried out to evaluate the laser velocimeter as a
velocity measuring technique in liquid-fuel turbulent flames and measurements of
mean and rms gas and droplet velocities were obtained for several test conditions.
The laser velocimeter used in these tests was identical to the one employed
in the gaseous fuel test program (Ref. 10). The measurements were made with a
dual beam velocimeter utilizing an argon ion laser and a crystal Bragg cell
which acted as a beam splitter and frequency shifted the first deflected beam. A
schematic of the optics and signal processing instrumentation is shown in Figs.
A-20 and A-21. The sensing volume determined by beam crossover volume, off-axis
collection and photomultiplier pin hole size was elliptic with principal axes of
0.2 mm and 2.0 mm, respectively. The velocity component sensed with this optical
arrangement lies in the plane of the two incident beams and is perpendicular to
their bisector. Single particle, time domain signal processing was used to build
up the velocity probability density distributions from which both the mean and
rms velocities were obtained using the following equations:
N .,
U=Z^L (A-2)
i = l N
cr =
N
In the present experiments, a minimum of 1000 instantaneous velocity determinations
was used to build up the probability densities. This number of determinations
results in a statistical error of less than 5 percent in the computer values of
both the mean and variance with a confidence level of 95 percent (Ref. 10). On
line signal processing to determine the local mean velocity, turb^ef.f^7
and probability density function was achieved using the instrumentation shown
87
-------
SCHEMATIC DIAGRAM OF THE LASER VELOCIMETER
FIG.A-20
TEST SECTION
16 BIT WORDS TO
MINI COMPUTER
88
-------
FIG.A-21
SCHEMATIC DIAGRAM OF THE LASER VELOCIMETER DATA PROCESSING EQUIPMENT
DATA SYSTEM
INTERFACE
MINI-COMPUTER
TELETYPE TERMINAL
CASSETTE TAPE
RECORDER
VISUAL DISPLAY OF DATA
16 BIT WORD FROM LASER VELOCIMETER
SIGNAL PROCESSOR (COUNTER)
DATA RATE 200 TO 40,000 WORDS PER SEC
76-03-270-9
89
-------
in Fig. A-21. A limited number of measurements of droplet velocity were made
using a frequency-tracker. These measurements were limited to those regions of
the burning spray where valid data rates were in excess of 20,000/sec.
The optical sensitivity of the forward scatter system used in the present
study was such that naturally occurring submicron particles could be used for
the gas velocity determinations in regions far from the fuel spray. However,
to increase the signal to noise ratio and thus, increase the data acquisition .
rate in these regions, the air flow was seeded with micron sized particles dis-
pensed from a fluidized bed. Since for these flows more than 95 percent of the
turbulence energy is likely to be in scales corresponding to Eulerian frequencies
below 25 kHz (turbulence scales less than 1/10 of the combustor dia), errors due
to particle response (in the Lagrangian frame) should be negligible (Ref. 10).
To determine gas velocity in regions of moderate fuel droplet concentration the
air flow was selectively seeded as discussed later.
To avoid problems associated with directional ambiguity (which can result
in data interpretation errors in highly turbulent and/or recirculating flows)
(Ref. 18), zero velocity frequency offset was achieved by combining the primary
and modulated beams at the detection volume where they generated moving fringes
so that a stationary particle produced a Doppler frequency, f . Thus, in the
flow field, moving particles generated Doppler frequencies of f jf fp depending
on their velocities normal to the moving fringes. Hence, the sign as well as
the magnitude of the instantaneous velocities could be determined as follows:
2 sin 6/2
where A is the wavelength of the laser light and 6 is the angle between the
incident laser beams.
However, in the case of fuel droplet velocity studies there is another
reason to use a moving fringe laser velocimeter system. Consider first a
stationary fringe system where the peak to peak fringe spacing, d, is
2 sme/2 a
when 6 is small.
Since the (1/e^) diameter of the focal region is given by
= X (A-6)
7T 2D
90
-------
where f is the focal length of the lens and 2b is the beam diameter at the
focusing lens. The number of fringes contained within the probe volume is
therefore, given by '
w 2b (A-7)
where D is the beam spacing at the focusing lens. Thus, the maximum number of
fringes is independent of the focal length of the focusing lens and is determined
solely by the ratio of input beam spacing to diameter ratio. Increasing the
number of fringes will increase the accuracy of the velocity measurements by
providing a greater number of signal cycles from which the average value may be
determined. Since most commercial counting devices require a minimum of eight
cycles this provides a lower fringe number limit. Unfortunately, most practical
liquid fuel sprays have a size distribution over a wide range (up to 250 vim)
with typical mean droplet sizes between 75 and 150 um. So this minimum fringe
number requirement may raise a problem of spatial resolution associated with
stationary fringe systems, since optimum signal/noise requirements dictate that
the fringe spacing should be greater than or equal to the maximum fuel droplet
diameter. Thus, for 250 um droplets, the focal volume diameter would be approxi-
mately 2 mm. Even with off-axis light collection the probe volume dimension
limiting spatial resolution (i.e., 2b0/sin6/2), which typically is an order of
magnitude greater than the focal volume diameter (£ 2 cm), would render the
spatial resolution of most optical systems open to question.
Fortunately, frequency biasing increases the number of effective fringes in
the focal volume and hence, the number of cycles of useful information. Again,
the fringe spacing is given by Eq. (A-5) but now the fringes are propagating
with a velocity U given by
thus, the actual number seen by a droplet is
Ns=N[f0/f0±l]
In the present system Ng <_ 45N so that sufficient fringe crossings can be
achieved with adequate spatial resolution.
In the experiments using iso-octane, measurements were made using the green
line (X = 514.5 nm) from the laser. Because of the intense lummosxty from the
91
-------
No. 2 fuel oil flames, an improved signal-to-noise ratio could be obtained using
the blue line (A = 488.0 nm). However, a limited number of measurements on No. 2
fuel oil flames were made using the green line to verify the velocity data.
Laser Velocimeter Statistical Errors and Particle Dynamics
Statistical confidence levels within stated error limits in the determination
of both the mean and variance of any quantity with a Gaussian probability variation
may be defined according to Ref. 19 as
C. I I w I \ r~ * ' / /IM / , - _»
(A-10)
Error=P/ISX2- cr2i\<
N-l
(A-ll)
where x is a random variable; N is the number of samples;
N
= calculated mean (A-12)
N
N LUI/"' ' J
Q2 ' TV fv -x)2l = calculated variance
sx --[1 Uj x, j
2
3 is the true mean and o is the true variance.
For the particular case of the laser velocimeter measurements let us replace
the random variable, x, by the measured Doppler frequency (fD). Then the confi-
dence level for the mean velocity determinations may be written
fn-/3B- Sf I (A-1A)
-------
since
'f .
u
(A-15)
we see that
< K.
(A-16)
The confidence level for the standard deviation may be written as
2 Sf-o-f
y
(A-17)
Now since sf and fD are functions of the same random variable (f ), the error in
au/U is the sum, not the square root of the sum of the squares of each error i e
the confidence level in au/U = YM + YT-
For normal distribution functions, confidence levels may be calculated using
the following table:
0.5
0.675
0.68
1.00
0.9
1.6
0.95
1.96
0.98
2.33
0.99
2.57
For example, if at a particular location, the local turbulence level was 1 percent
and 100 instantaneous velocities were measured, 50 percent of the mean measurements
would be in error by less than 0.0675 percent of the true value. Whereas only 1
Point in a hundred would be in error by more than 0.257 percent.
For a spherical particle of diameter D suspended in a sinusoidally vibrated
column of air and acted on by Stokes drag, the ratio of particle velocity to
gas velocity can be expressed as (Ref. 20):
93
-------
2.- 1/2
where
a = d. /1 +
where u and u are the rms velocities of the particle and the gas p is the
particle density, f is the vibration frequency, £ is the molecular mean free path
of the gas and K is the Cunningham constant (£ 1.8 for air). Thus, a 5 urn phenolic
resin microballoon in air at ambient conditions will follow velocity fluctuations
up to 10 kHz within 10 percent. Power spectral density measurements in the shear
layer of nonreacting jets (Ref . 20) indicate that for the reacting flows investi-
gated in the present study more than 95 percent of the turbulence energy will be
associated with Eulerian frequencies below 25 kHz. Hence, the scale of the
smallest energy containing eddy will be on the order of
\-. u/f ~ ioOm/sec/25 kHz = 4xiO"3m (A-20)
In the Lagrangian frame, this scale corresponds to the frequency on the order of
f - (U- u C)/X ~ 20m/ sec/4 xiO'3m= 5kHz
so that errors due to particle response should be negligible.
(A-21)
94
-------
APPENDIX B
PHASE-DISCRIMINATING PROBE TEST PROCEDURES
Hydrocarbon Data
For each mapping test condition, a series of measurements was made by traver-
sing the phase-discriminating sampling probe at various axial locations to obtain
percent carbon as a function of radial distance. Before and after each test the
calibration of the hydrocarbon analyzer was checked by using three different
calibration gas compositions (100 percent, 4.5 percent and 0.09 percent methane).
After the combustor airflow conditions were established, the probe was traversed
to the desired radial position and the sampling line valves opened. Isokinetic
conditions were established in the total sample line at the probe tip by adjust-
ing the total sample flow rate to match the probe and combustor static pressures.
The vapor-sample mass flux was set to between 10 and 20 percent of the total
sample mass flux, and the suction flow rate was set to between 30 and 60 percent
of the total-sample mass flux. The background level of hydrocarbons in the
combustor flow was then measured to insure that no hydrocarbon contaminants were
present. Combustor fuel flow was initiated, the flow was ignited, and steady
state conditions were established within the combustor. Sample line flow rates
were readjusted and the attenuation levels on the electrometer were adjusted
to give reasonable concentration peak heights on the chart recorder. For most
tests, two measurements of the hydrocarbon content of the total sample and two
measurements of the vapor sample were obtained at each probe condition.
For each of the flow field mapping tests, see Table B-l, the phase discrimin-
ating probe was positioned at window port locations 3 through 6. The axial location
of the probe tip at these four positions is given in Table B-l. For each sample,
values of the integral of percent carbon versus time were obtained from the
mechanical integrator output. This information was combined with the hydrocarbon
analyzer calibration curve to obtain the local value of percent carbon. Tabulations
of the experimental data are presented in Appendix H. The tables indicate the
combustor test conditions and present data on the percent carbon in the total
and in the vapor samples at various probe psotions.
To assist in interpretation of the experimental results, the data were re-
duced to isopleth form. To accomplish this, plots of percent carbon as a function
of radial position were generated for each of the four axial posxtions (e.g.,
Fig. B-l), and data crossplotted to obtain the isopleths.
95
-------
TABLE B-1: PHASE-DISCRIMINATING PROBE TEST MATRIX
TEST
1
4:
5
6
8
11
FUEL
ISO-OCTANE
ISO-OCTANE
ISO-OCTANE
NO. 2 FUEL OIL
NO. 2 FUEL OIL
PROPANE
PRESSURE ATM
1
3.3
1
1
1
1
SWIRL
0.3
0.3
0.6
0.3
0.3
0.3
TEMP °K
533
533
533
533
755
533
AXIAL WINDOW
3
X
X
X
X
X
4
X
X
x
X
X
5
X
X
X
X
X
6
X
X
X
X
X
PHASE-DISCRIMINATING PROBE TIP POSITION
0=0.65
WINDOW
3
4
5
6
PRiOBiE
TIP LOCATION,*
X(cm) X/D
14.83
31.83
48.82
65.81
1.21
2.60
3.99
5.38
"FUEL INJECTOR TIP is AT x = o
D = 12.23 CM
76-06-145-1
96
-------
TYPICAL RADi AL PROFILES OF PERCENT CARBON
ISO-OCTANE
TEMP = 533 °K
PRESS = 1.0 ATM
S = 0.3
FIG. B-1
O TOTAL
X/D = 1.21
X/D = 2.6O
X/D = 3.99
X/D= 5.38
VAPOR
1.0
0.1
0.01
COMBUSTOR
; WALLS
RADIAL POSITION,
-1.0
76-09-2-9
97
-------
APPENDIX C
LASER VELOCIMETRY TEST PROCEDURES
The use of LDV in liquid-fuel combustor experiments presents unique problems
because of the presence of the fuel droplets. The basic mechanism for LDV
techniques is scattering of incident laser light by particles. If motion of the
spray were the only phenomenon of interest, the potential problems would be
limited to the effects of particles outside of the measuring volume on the
signals. These effects (opaqueness, background luminosity, beam wander and
divergence) were not apparent. Thorough investigation of combustor flows, how-
ever, requires knowledge of the gas velocities as well. To obtain such information,
the air supply was seeded with particles. Two methods of data analysis were used
to distinguish between gas and spray velocities. In regions where the spray drop-
let density was relatively low, testing with and without seeding in the air supply
allowed the two velocities to be separated. In regions of high droplet density,
a method based on relative signal (or visibility) was employed. Selective seeding
of the air supply was used in regions of the flow where the gas and fuel droplet
velocities were substantially different so that characteristic bi-modal velocity
probability density functions were apparent, as illustrated in Fig. C-l. With
seeding, the probability density function at R/R0 = 0.75 is heavily weighted toward
the local gas velocity since the seen particle number density is much larger than
the number of fuel droplets in this region. The situation is reversed as the point
of measurement approaches the mean droplet spray trajectory (R/RQ = 0.63). The
method used to separate the two mode velocities is illustrated in Fig. C-2, where
probability densities obtained with and without airstream seeding are presented.
Without seeding, the fuel droplet velocity distribution is determined directly
since the natural particle data rate is small. This distribution is normalized
by the total number of velocity determinations and then subtracted from the bi-
modal distribution obtained at the same location with air seeding. The result
represents the local gas velocity distribution from which the mean velocity and
the variance can be determined.
Typical mean axial velocity profiles obtained with and without seeding are
presented in Figs. C-3, C-4, and C-5. Close to the injector (X/D = 0.164) large
differences between local fuel droplet and gas velocities occur and there is
evidence of gas flow recirculation which is indicated by significant numbers of
negative seed particle velocity occurrences close to the half radius locations.
At X/D = -.409 there is agreement between velocity measurements obtained with and
without seeding which could be interpreted to mean that, except in the wall region,
the fuel droplets are following the local gas flow. However, away from the wall
the velocity data acquisition rates showed no significant changes when seed parti-
cles were introduced; thus, it was apparent that fuel droplets were dominating
98
-------
FIG. C-1
AXIAL VELOCITY PROBABILITY DENSITY FUNCTIONS WITH SEEDING
ISO-OCTANE/AIR
SWIRL = 0.6
1 ATM
TAlR = 533°K
0 = 0.65
X/D = 0.164
0
-20
0 20 40
INSTANTANEOUS VELOCITY (M/S)
60
76-10-29-12
99
-------
FIG. C-2
AXIAL VELOCITY PROBABILITY DENSITY FUNCTIONS WITH AND WITHOUT SEEDING
ISO-OCTANE/AIR
SWIRL =0.6
1 ATM
TA(R=533°K
0=0,65
X/D = 0.164
100 -
CO
LLI
O
z
01
DC
0
O
O
LL
O
DC
LLI
03
0 10 20 30 40 50
INSTANTANEOUS VELOCITY (M/S)
76-10-29-1
100
-------
FIG. C-3
MEAN AXIAL VELOCITY PROFILE
ISO-OCTANE/AIR
SWIRL=0.3
1 ATM
'AIR
= 533°K
0=0.65
X/D=0.164
AXIAL VELOCITY (M/SEC) -r 40
A SEED
A NO SEED
RADIAL POSITION, R/R0
76-10-29-5
101
-------
FIG. C-4
MEAN AXIAL VELOCITY PROFILE
ISO-OCTANE/AIR
SWIRL = 0.3
1 ATM
TA|R=533°K
0=0.65
X/D =0.409
30
O
S 20
O
O
'
<
0
-1.0
SEED
Q NO SEED
-0.5
0.5
1.0
RADIAL POSITION, R/R0
76-10-29-7
102
-------
FIG. C-5
MEAN AXIAL VELOCITY PROFILE
ISO-OCTANE/AIR
SWIRL = 0.3
1 ATM
TA|R = 533°K
0 = 0.65
X/D=1.64
40
o
LU
tn
o
g
LU
X
<
30
20
10
-1.0
SEED
O NO SEED
_L
-0.5 0 0.5
RADIAL POSITION, R/R0
76-20-29-6
103
-------
the velocity probability density distributions at this location. Except in the
region close to the wall where a sufficiently high seed particle/fuel droplet
number density ratio could be achieved, there was no evidence of bi-modal dis-
tributions and therefore gas velocity data could not be extracted. Farther down-
stream (X/D = 1.64) droplet concentrations are sufficiently reduced by evaporation
and combustion so that the addition of seed material can once again influence
the data acquisition rate. Thus, the agreement between the seeded and unseeded
velocity measurements at this location shows that the fuel droplet and/or particu-
late combustion products are convected with the local gas velocity.
In regions of high droplet number density (X/D = 0.409 for example) an alter-
nate method which utilizes particle visibility variations can be employed to in-
fer local gas velocity. The visibility of the photomultiplier output signal which
is a measure of the relative amount of oscillation in the signal, is defined for
the current application as
v (c-i)
max min
where I and I . are the collected scattered Intensities when a particle is
max min
centered in a bright and dark fringe, respectively.
The visibility is dependent on several factors: incident laser power, measur-
ing volume dimensions, location within measuring volume of the particle, and
electronics conversion factors. Incident laser power was monitored on the laser
power supply and held constant. Fixing the incident and receiving optics for the
entire test, the measuring volume dimensions were held constant, subject to
beam wander due to temperature and spray effects on local index of refraction.
This latter effect was deemed small, principally because of no noticable dropout
which would occur if the volume produced by the incident beams did not coincide
with the collecting volume. This is not a completely satisfactory argument, how-
ever, since the incident volume was necessarily large because of the small included
angle between the incident beams. It is conceivable that the receiving volume
could move relative to the incident volume producing apparent signal variations
while maintaining essentially continuous signal output. The location within the
measuring volume (defined by the overlap of the incident and receiving volumes)
determines the incident power level since the Guassian intensity distribution of the
incident beams results in an intensity distribution within the volume. This problem
can be minimized by controlling the offset frequency. Careful selection of the value
can limit the fraction of the incident volume over which the residence time of
particles within the volume is sufficient to permit the counter processors to validate
the data. Again, the relatively small included angle between the beams limits the
effectiveness of this procedure. The electronics conversion factors convert a scat-
tered light intensity to a signal voltage. Within the test period for the current
effort, variations in these factors were negligible.
104
-------
The use of the visibility, subject to the restrictions of the previous para
graph, is dependent upon the interaction of particles and gas in the following
manner. For particles in an airstream, the velocity lag of the particle relative
to the gas is a function of the drag on the particle. For simple models the
velocity lag is a function of the square of the particle diameter. For a limited
range of particles, assumed spherical, the signal level or visibility can be
considered linear with particle size. (Acutal Mie scattering is much more com-
plicated.) The result is a dependence of velocity lag on the square of the visi-
bility defined by Eq. C-l. An example is shown in Fig. C-6, where extrapolation
to zero visibility coincides with the gas velocity (anticipated in this instance
by the secondary peak noted on the histogram shown schematically) . Visibility
variations were achieved by altering the trigger level necessary for a signal
from a particle to be input to the processor. Thus it is often possible to distin-
guish between spray and gas velocities in this manner.
Careful consideration should be given, however, to the various limitations
and assumptions necessary for the procedure. For example, well inside the fuel
spray, the variation of velocity with particle size was found to be the opposite
of that for Fig. C-6. It is possible that this apparent lower gas velocity com-
pared to the spray velocity is indicative of a gas recirculation region which has
little influence on the spray with its relatively higher momentum. It is also
possble that the mechanism is not applicable for this case. If the particle
momentum is such that the spray droplets are not carried by the gas flow (with a
velocity lag due to particle drag) but rather interact in some more complicated
manner, the general dependence of velocity on visibility may change over the range
tested.
In conclusion, therefore, the use of visibility as defined within this report,
provides additional information subject to several restrictions and assumptions.
The general application of the method would require more detailed attention to
these procedural limitations; but in principle, the potential exists to not only
extract local gas velocity but possibly also relative particle size distributions.
This ultimate effect of the limitations on the accuracy of the results is unknown
based on work to date.
A new method was used to determine droplet spray trajectory on the present
work which provides a relatively simple and convenient means of trajectory docu-
mentation. The technique, which involves droplet velocity number density
measurements, is illustrated in Fig. C-7 which shows the velocity data rate
variation across the combustor at X/D = 0.164. The two peaks define the time-
averaged hollow cone spray location. Measurements at a series of axial stations
can then be used to determine the time-averaged spray trajectory.
The overall root-mean-square turbulent velocity measurement techniques applied
in the program represent the simplest characterization of combustor turbulent flow
field. Such measurements are therefore limited in the extent to whxch
.
used to- obtain physical insight into the mixing process. ^Tm*t°*
velocity length scales as a function of frequency and position would be desirable
105
-------
FIG. C-&
INFLUENCE OF VISIBILITY ON MEASURED LOCAL MEAN VELOCITY
100
in
o
2
111
CC.
13
O
O
O
50
20
40
M/S
20
o
LU
in
o
O
X
<
15
10
10
(VISIBILITY)2 (VOLTS2)
76-10-29-3
106
-------
FIG. C-7
DROPLET NUMBER DENSITY PROFILE
ISO-OCTANE/AIR
SWIRL = 0.6
1 ATM
$=0.65
X/D = 0.164
400 -
300 -
CJ
LU
V)
LU
DC
<
<
Q
200 -
100 -
RADIAL LOCATION ,R/R0
76-10-29-8
107
-------
since they would assist in the understanding of the relative importance of varying
length scales on the mixing process. For example, is the mixing process governed
by the large eddy structure or do such eddies act as a gross transport mechanism
with the mixing rate determined by small scale eddy interaction? In general,
LV measurement of turbulent velocity cross-power spectral density, which would
provide frequency dependent length scale information, does not appear practicable
in combustor flows at this time. Two-point LV velocity correlations are possible,
however, and these measurements would provide eddy convection velocity information
and frequency integrated (overall) eddy length scales. From the behavior of the
cross-correlation function sin space and time (i.e., zero time delay amplitude
verses separation distance and amplitude versus delay time at fixed separation
distance) it may be possible to obtain information concerning the various length
scales and their associated convection velocities. Such two-point measurements
are to be preferred to single point correlations (auto correlations) for which
Taylor's hypothesis ("frozen-flow" assumption) must be invoked to infer length
scales from Eulerian time scales. For the high turbulence levels encountered in
a combustor, Taylor's hypothesis would be expected to be invalid.
Prior to performing such two-point LV correlations in a combusting flow,
validation of the technique should be carried out in a cold flow simulation. In
addition to establishing the ability to perform correlations when the data rate
is insufficient to permit use of a tracker, such a simulation would provide
correlation data of use in the planning of a combusting flow measurement program.
108
-------
APPENDIX D
LASER HOLOGRAPHY TEST PROCEDURES
In the tests which employed the holographic system, the optical apparatus was
aligned with the first window in the combustion rig so that the field of view
included the region immediately downstream from the fuel injector, Fig. 1. A
cylindrical ring containing calibration wire and having an internal diameter of 5 cm
was installed in the window so that the unobstructed viewing area was defined by
a 5-cm dia circle with a center located 4.2 cm downstream of the injector tip.
Hence, the nozzle tip was outside the field-of-view.
During initial testing of the holographic system an interference filter,
which was mounted adjacent to the quartz window in the test apparatus, failed due
to overheating. To preclude further failures, the filter-holder was moved away
from the window, cooling air was flowed over the surface of the filter, and the
combustor was operated only for the time required to establish steady flow conditions.
To obtain the holographic data, the optical components were aligned and the
magazine containing eight unexposed film plates was loaded. Airflow was initiated,
the electric air heater was activated and the combustor was brought up to tempera-
ture as evidenced by a steady reading of the inlet air temperature. Fuel flow was
initiated and a hologram was exposed to record the behavior of the spray in non-
combusting flow. The duration of the fuel-on portion of this test was less than
ten seconds. Holograms were not recorded during the first five seconds to insure
that a steady fuel flow had been achieved prior to the recording of data. After
fuel shut-off, hot air continued to flow for approximately one minute to purge
the combustor of any accumulated fuel. The film magazine was advanced, fuel flow
was initiated, the flow was ignited and after a period of approximately ten seconds
a hologram was exposed to record the spray characteristics in combusting flow.
The combustor was shut down and the interference filter was replaced with the
gelatin filter and the test sequence repeated.
After development of the holographic plates, the reconstructed images of the
holograms were examined and the better of the two holograms taken at each condition
(one with the gelatin filter, one with the interference filter) was selected for
detailed examination. The first step in the reconstruction process was alignment
of the holographic plate relative to the incident reference beam. The tripod-
mounted loupe was then placed so that the entire f ield-of-view could be examined
by moving the rack-and^pinion traversing mechanisms. The surface coordinates
of the hollow cone spray were determined by focusing on the far and near surfaces
as illustrated in Fig. D-l. This technique was used rather than focusing on the
upper and lower edges of the cone because those edges were generally outside the
viewing area. The spray surface is,in fact, a diffuse layer of droplets of finite
109
-------
FIG. D-1
METHOD USED TO ESTABLISH SPRAY SURFACE COORDINATES
TYPICAL DROPLET
SIZE SURVEY REGIONS
FUEL INJECTOR
TYPICAL SPRAY
SURFACE DATA
POINTS
FLOW DIRECTION
76-09-2-1
110
-------
FIG. D-2
TYPICAL SPRAY COORDINATE DATA
ISO-OCTANE, NO COMBUSTION
T = 541 °K
S = 0.6
P = 1 ATM
DERIVED CONE ANGLE = 70.5 DEC
-5.0
2.0
4.0 6.0 8.0
DOWNSTREAM DISTANCE, z (CM)
10.0
76-09-2-6
111
-------
thickness; the "surface" coordinates recorded are the coordinates of the center of
the region being viewed. The center was established by fixing the x and z
coordinates and traversing along the y coordinate through the volume containing
droplets. The image viewed was of a wave of droplets passing across the object
field. The y coordinate was established by centering this wave in the object
field.
Estimates of the mean droplet size in the spray were obtained by counting
droplets of various diameters at several locations as shown in Fig. D-l. These
regions were established at values of z in the range from 2.5 to 4.7 cm. The value
of z could not be fixed at a specific value for all holograms because the location
of the injector tip, which was not in the field of view, could only be estimated
at the time of data acquisition. The volume of the individual regions examined
for droplet size were defined by a 3.6 x 1.0 mm rectangle and a depth equal
to the spray layer thickness. Typically, seven volumes were examined in each
hologram.
Tabulated results giving the spray surface coordinates and the derived spray
cone angle are presented in Appendix M. The spray coordinates given are for a
transformed coordinate system in which the z axis lies along the centerline of
the injector and the apex of the cone lies at the injector tip. The reported cone
angle was obtained by determining the best fit between the data and the surface
coordinate of a right circular cone by use of the regression analysis described
in Appendix K. The estimated error in the derived cone angle was approximately
2 to 3 deg for most of the cases examined. A plot showing a typical distribution
of data as projected on the y-z plane is shown in Fig. D-2. Data were not obtained
in the x = 0 plane. The projected edges of the cone appear as boundaries for the
data.
Measurements of the thickness of the spray cone and of the size of the
observed droplets downstream of the injector tip also are given in Appendix M.
The coordinates y-i and y^ represent the observed coordinates of .the inner and
outer edges of the spray on the side of the hollow-cone spray closest to the
observer. The coordinate z gives the downstream distrance at which the measure-
ments were obtained. Occasionally a droplet might be observed outside of these
coordinates; however, it is estimated that over 90 percent of the observed drop-
lets reside between yi and J2 i-n anY °f tne measurements. Because of the limited
number of measurements made on spray thickness, the error in the spray thickness
determined from the holograms is estimated to be approximately 30 percent. How-
ever, there is generally good agreement between the estimates of spray thickness
obtained from the laser velocimeter data and the values obtained from the holograms.
The droplet count tables give the number of droplets in a range of sizes from
2 mils to 10 mils in diameter. Conversion of the data into microns and application
of the correction for the magnification caused by the fact that the holograms were
reconstructed using a laser wavelength, different from the wavelength used to
generate the hologram (by a factor of approximately 1.1) results in the numbers
shown.
112
-------
Droplet size could be measured with greater accuracy than required by the fineness
of this classification schedule, however, a finer classification was not warranted
for the study.
The Sauter mean diameter (SMD) generally is regarded as the appropriate
diameter to be used to characterize a spray fqr use in analyses in which surface
phenomena such as vaporization or combustion are of interest. The SMD is computed
by the expression:
Z-3
n-,Ui
SMD -
(D-l)
where:
n* - number of droplets of diameter d.
d^ - droplet diameter
The SMD ranged from 100 to 150 microns for the holograms examined. In the table
of reduced data, the quantity N-counted, which is determined by dividing the
total mass of the counted droplets by the mass of the droplets having a diameter
equal to the SMD is given. This quantity represents, on average, the number of
droplets counted and is approximately 50 for most of the holograms examined. The
quantity N-measured represents the number of droplets passing the plane per unit
time at which the droplet measurements were made based on the number of droplets
counted, the measured spray diameter, and an assumed value for the velocity of the
droplets.
N- measured = N counted x x
b
^«Ds-H)2 - (ps- t)?)xb
axbxNblock
where;
N
D -
t -
a, b
, ,
block
V
Spray diameter
Spray thickness
Dimension of block
Number of blocks counted
Droplet velocity
The magnitude of the droplet velocity, V, was taken to be 19 m/sec, a value
obtained from LDV measurements of the droplet velocity in an iso-octane spray
at a location five centimeters downstream of the injector tip (Fig. C-4, Appendix
C). The quantity N-calculated represents the number of droplets per unit time
passing the plane as computed from the measured fuel flow and the derived mean
droplet size;
113
-------
N- calculated-- - (D-3)
-
where Wf _ Fuel flow
p - Fuel density
dm - Mean droplet diameter
In general it was found that the value of N-calculated was less than that of
N-measured.
The value of the fuel flow rate as calculated from the measured droplet size,
number density and velocity (as determined from the laser velocimeter measurements)
was generally found to be one half the measured fuel flow rate. This discrepancy
could arise from an under estimate of the mean droplet size, failure to account
for the mass of the small droplets in the spray (which are not counted), the
influence of vaporization or an under estimate of the velocity of the droplets.
The fact that little change in the magnitude of the discrepancy occurred when
changing from combusting to noncombusting flows suggests that the vaporization
effects are not the cause. The laser velocimetry data indicates that it is unlikely
that the spray velocity is significantly greater than the assumed value of 20
m/sec. For typical droplet size distributions obtained for pressure atomizing
nozzles, droplets smaller than the SMD can account for as much as 20 percent of
the mass flow; therefore, this factor could account for a portion of the discrepancy.
Finally, because the cube of the SMD enters the calculations, an error of 17 percent
in the droplet sizing measurements could result in the observed difference. It
is probable that the latter two factors were major contributors to the observed
difference.
Several observations of general interest were made during examination of the
reconstructed holograms. First, no direct evidence of combustion could be observed
in any of the holograms; i.e., there was no change in resolution or evidence of
index of refraction gradients for the cases for which combustion was occurring.
During testing, there was a large amount of radiation emitted from the gases
in the region being examined by the holographic system, therefore combustion
gases were certainly present in this region, at least on the time average.
The quality of the holograms was more or less independent of the type of
filter used to block visible radiation from the combustor. It did not appear that
the superior ability of the interference filter to block transmission of the
combustion-generated radiation had a beneficial effect on hologram resolution,
nor did the optical qualities of the filters appear to affect resolution (as
determined from the noncumbusting flow cases). In each of the sprays, there
were a few large droplets having diameters approximately 5 times the mean droplet
114
-------
diameter, and the size of these droplets appeared to increase with distance from
the injector. These observations suggest the occurrence of agglomeration of
mailer droplets; however, there are insufficient data to substantiate this point.
-------
APPENDIX E
TYPICAL PROPERTIES OF LIQUID FUELS
Specific gravity (@ 298 K)
H/C
Wgt. fraction N
Viscosity (CS @ 298 K)
Surface Tension (dynes/cm
@ 298 K)
Initial Boiling Pt. (deg K)
End Boiling Pt. (deg K)
Specific Heat (Cal/gm K)
Heating Value (Cal/gm)
Heat of Vaporization
(Cal/gm)
Molecular Wgt.
Stoichiometric Ratio Obywgt.) .065+
Propane Iso-Octane
0.70
2.67 2.25
0 0
0.8
18
231 373
-
0.52
1106U 10600
102 65
1+1+ lll+
) .061+ .066
No. 2 Oil
0.85
1.75
v
5-20 x 10-6
5.3
21+
1+5U
603
0.1+5
10570
153
195 (nominal)
.069
116
-------
APPENDIX F
COMBUSTOR HEAT BALANCE
A thermal balance was performed on the combustion system for a typical operating
ndition to evaluate the magnitude of the heat transferred to the combustor walls
nd to provide a check on the accuracy of the exhaust gas temperature measurements.
The heat transferred from the system was determined from measurements of the flow-
rate and the temperature rise of the cooling water, The results of the heat
balance are summarized below:
QIN
Swirl No. =0.3
p = 1 atm
T = 533UK
air
Air Heater
Fuel:
Iso-octane - Air
Equivalence Ratio = 0.65
m . = 0.137 kg/sec
air
7.7 kg cal/sec
61.9 kg cal/sec
69.6 kg cal/sec
QOUT
Combustion Products:
Cooling Water:
Unreacted Fuel:
43.04 kg cal/sec
17.90 kg cal/sec
0.72 kg ca/sec
61.6 kg cal/sec
117
-------
APPENDIX G
TABUIATED TEMPERATURE DATA
118
-------
TABLE G-fl. TEMPERATURE DISTRIBUTIONS - TEST NO. 1
Iso-Octane Inlet Air Swirl No. = 0.3
$ = 0.65 - .01 Inlet Air Temperature = 532- 5° K
Pressure = 1.0 atm
oh Temperature, °K
' o
-0.88
-0.75
-0.67
-0.62
-0.50
-0.37
-0.25
-0.12
0.00
0.13
0.25
0.38
Q.ll-2
0.50
0.58
0.63
0.67
0.75
0.83
0.87
0.92
1.00
X/D=0.3U
KM
__
MM
MM
MM
MM
MM
__
__
568
677
692
630
555
516
U76
Ull
1.73
1109
1^31
--
1622
1706
1718
1662
16U6
1690
1787
1701
1585
__
13^7
1112
<9lH
7^3
1466
3.12
133U
1558
1718
1830
186U
1779
175)4
1888
1870
1776
1608
~
1^3*4
1233
1056
858
T3U
h.52
135^
1557
~
16^8
1769
1887
1868
1796
1910
1906
1809
1656
--
1535
13^8
1107
866
1426
5.91
1265
1^31
1582
1708
1815
1761
1721
1800
18^7
1717
1587
MM
1^55
1310
113^
886
U58
1^.59
1381
lUlU
1)458
1521
1558
1573
1550
1508
1)459
1392
1327
1276
1210
1105
5)4-0
1X9
-------
TABLE G-2. TEMPERATURE DISTRIBUTIONS - TEST NO.
Iso-Octane
$ = 0.65 ± .01
Pressure =3.3 atm
Inlet Air Swirl No. = 0.3
Inlet Air Temperature = 529 ± 7° K
R/Rr
Temperature, K
-0.89
-0.75
-0.68
-0.62
-0.50
-O.U5
-0.37
-0.25
-0.12
-o.oU
0.00
0.13
0.17
0.25
0.38
0.50
0.58
0.63
0.75
0.79
0.87
1.00
X/D=1.73
1105
1^9
1066
.
-
836
771
702
681
690
1163
12lU
1058
725
613
582
3.12
928
1^05
-.-
__
--
876
810
786
802
1382
1636
Ih07
1055
857
769
678
U.52
1197
1^53
__
__
__
--
888
__
860
__
' 1075
1180
--
1395
1529
1*4-05
--
10U7
8lO
--
713
6Ul
5.91
753
1133
__
__
1291
__
__
877
787
-_
779
1155
--
1^405
158U
1502
__
1270
990
__
77^
660
1^.59
M»
-
1192
__
_
1269
__
880
MM
1107
__
__
1296
123^
1191
_«
1095
*«
»M
983
638
120
-------
TABLE G-U TEMPERATURE DISTRIBUTIONS - TEST WO. 5
Iso-Octane Inlet Air Swirl Ho. = 0.6
§ = 0.65 Inlet Air Temperature = 530 - 50
Pressure = 1.0 atm
Temperature, °K
-0.89
-0.75
-0.6?
-0.62
-0.58
-0.50
-0.37
-0.35
-0.25
-0.18
-0.12
0.00
0.13
0.25
0.31
0.38
0.50
0.63
0.75
0.8l
0.87
1.00
X/D=1.73
1U88
1797
__
1766
15^5
1085
11U7
1^97
1639
-i-'--' ^j ^
1801
17^7
1^-78
1159
72U
393
3.12
15^3
1908
__
1880
1678
«M
1122
1^76
l6lO
1736
1869
1931
1632
1256
893
738
385
U.52
1168
17^5
1858
__
1169
_-
1070
__
1339
1360
1605
19^0
«
1898
lU79
1090
373
5.91
1255
1679
1809
1775
1821
785
797
902
1101
12Ul
1376
1832
--
1673
122U
897
718
6UO
U29
1^.59
_-.
10^6
923
10U5
1225
1597
--
1375
1225
1129
1087
^
121
-------
TABLE G-V- TEMPERATURE DISTRIBUTIONS - TEST NO. 6
No. 2 Oil Inlet Air Swirl No. = 0.3
§ = 0.6? Inlet Air Temperature = 529 ± 6° K
Pressure = 1.0 atm
R/R Temperature, °K
X/D=1.73 3.12 If. 52 5.91 lU.59
-0.89 1088 1*4-52 15*16 1520
-0.75 12lU 1523 16^5 l6ll
-0.66 ~ 11+39
-0.62 131*3 1605 1711 1662
-0.50 11+51 i681+ 1738 1719 1501
-0.37 1^72 1726 1803 1771
-0.33 - - - 1538
-0.25 3M7 1717 176U 1789
-0.17 - - ~ - 1553
-0.12 iU2U 1669 1711 1763
0.00 11+28 1720 1736 1761+ 15U2
0.03 -- - 1711
0.13 1^06 1723 17^5
0.17 -- -
0.25 129^ 1639 1693
0.33 -
0.38 1195 1585 1653 1595
0.50 1081 1525 1579 1551
0.63 1008 ih66 1502 11+89
0.67 ~ __ 1267
0.75 92U ihoo 138^ iUio
0.83 ~ - - - 1206
0.87 837 1290 1211 1276
l.OO 623 685 739 762
122
-------
TABLE G-5 TEMPERATURE DISTRIBUTIONS - TEST HO. 8
Wo. 2 Oil Inlet Air Swirl No. = 0.3
$ = 0.6? Inlet Air Temperature = 751 ± k° K
Pressure = 1.0 atm
Temperature, °K
-0.89
-0.75
-0.68
-0.62
-0.50
-0.37
-0.33
-0.25
-0.16
-0.12
0.00
0.13
0.17
0.25
0.33
0.38
0.50
0.63
jf
0.67
0.75
0.83
0.87
1.00
X/D=1.73
lUUo
1^78
_«
152U
1592
1581
__
1559
1536
1553
1550
1U97
1^21
1336
13^1
M-B
1308
1050
3.12
1667
1797
_-
1865
1918
1918
1883
__
l8U2
1903
1903
1857
1783
1738
1639
"""
155^
iSs
-.*>
l6U9
1736
1819
1899
19^2
1863
18U6
1905
1926
1831
1752
1671
158U
1^55
""""
1267
687
5.91
16^1-6
1693
1782
1890
1925
1872
1876
1879
1817
1716
1659
1599
1553
__
1303
695
1^.59
--
1^89
16U3
1721
1732
1698
159^
1^89
1383
1338
1235
613
123
-------
TABLE G-6. TEMPERATURE DISTRIBUTIONS - TEST NO. 11
Propane Inlet Air Swirl Ko. = 0.3
$ = 0.65 i .01 Inlet Air Temperature = 536 ± 6° K
Pressure = 1.0 atm
R/R0
-0.88
-0.75
-0.68
-0.62
-0.50
-O.U5
-0.37
-0.25
-0.16
-0.12
-o.ok
0.00
0.13
0.17
0.25
0.33
0.38
0.50
0.63
0.67
0.75
0.83
0.87
1.00
Temperature, °K
X/D=1.73
33k
1195
1398
1310
1193
1127
1118
__
1076
1129
1232
1379
132k
997
677
k97
375
3.12
1161
1367
1030
--
996
1088
__
1265
1387
MM
156l
1^77
1088
720
55k
__
U92
39^
^
102.k
1392
-_
lU?i
__
865
--
932
__
-_
I_o8k
1311
~
1^91
1529
989
__
570
--
1+76
370
5.91
868
lU2U
__
793
__
759
8oU
890
__
1019
M(W
1128
1239
__
15^5
__
i6k3
ik99
1095
MM
693
M*.
521
33^
1^.59
^ ^
MM
11+67
--
__
1580
1658
_
1628
1696
__
1691
__
1533
1396
1260
_
1098
__
l+8i
124
-------
APPENDIX H
TABULATED PHASE-DISCRIMIWATIWG PROBE DATA
125
-------
TABLE H-l
PERCENT CARBON DISTRIBUTIONS AS DETERMINED BY
PHASE-DISCRIMINATING PROBE
TEST NO. 1
Iso-Octane
$ = 0.65
Pressure = 1 atm
X/D = 1.21
R/Ro
.672
.446
.161
-.048
-.257
-.297
-.505
-.505
-.795
X/D = 2.60
.673
-457
.148
.141
-.053
-.259
-.508
Inlet
Swirl No.
= 0.3
Inlet Air Temperature =
D = 12
H0 = 6
.23 cm
.17 cm
533°K
Percent Carbon
THC
1.04
1.14
3.63
5.95
3-93
1.59
1.48
1.10
2.16
.917
.410
1.36
1.50
2.07
1.105
.29
VHC
1.15
1.28
4.67
3.48
2.15
2.26
1.14
1.89
.868
.388
1.47
1.44
l.4o
.740
.39
THC
1.10
1.17
3-53
6.41
3.58
1.42
1.06
2.48
.923
.371
1.63
1.24
.271
VHC
1.19
1.17
4.86
38.02
2.34
1.21
1.05
1.94
.860
-387
1.59
.771
.369
THC
1.08
1.16
5.54
1.39
2.27
1.96
1.15
VHC
4.60
1.17
1.14
1.58
126
-------
TABLE H-l (Cont'd)
PERCENT CARBON DISTRIBUTIONS AS DETERMINED BY
PHASE-DISCRMINATING PROBE
Iso-Octane Inlet Swirl No. = 0,3
I = 0.65 Inlet Air Temperature
D = 12.23 cm
533°K
Pressure = 1 atm
RQ = 6.17 cm
X/D = 3-99
R/R0
.593
.1*75
.222
-.023
-.318
-.550
-.785
X/D = 5-38
.675
.572
.1*51*
.336
.078
-.026
-.319
-.360
-.53^
-.672
-.779
THC
.561*
.127
.537
1.5
.296
.123
.985
.1*71
-191
.061
.105
.722
1.08
.01*05
-0639
.3^2
.685
VHC
.376
.106
.1*31*
1.31
.21*7
.08^*3
.670
.593
.0868
.129
.793
.607
.0305
.0317
.0551
.188
.1*62
Percent
THC
.5*
.ill*
.553
1.1*1*
.382
.083!*
1.02
.292
.0527
.111
.862
l.Ol*
.0536
.07
.3*9
.61*5
Carbon
VHC
-1*13
.0868
.532
1.32
.266
.103
.787
.556
.0871
.81*7
.658
.03^*9
.0596
.188
.1*1*7
THC VHC
.528 .396
.513
.235
.109
.766
.535
.071*1
.031*
.0762
127
-------
TABLE H-l (Cont'd)
PERCENT CARBON DISTRIBUTIONS AS DETERMINED BY
PHASE-DISCRIMINATING PROBE
Iso-Octane
$ = 0.65
Pressure = 1 atm
Inlet Swirl No. = 0.3
Inlet Air Temperature = 533°K
D = 12.23 cm
RQ = 6.17 cm
X/D = 5.38
R/Ro
.560
on ]i
* OU-
.021
-021
-.206
-.599
THC
.013
.013
.101+
.11+
.113
.36
2.25
VHC
.0155
.02
.056
.086
.083
.3
1.2
Percent
THC
.0077
.0059
.093
.11
.36
1.58
Carbon
VHC
.0275
.011+
.058
.073
.081
.3
1.25
THC VHC
.0187 .0235
.16 .081
1.81
128
-------
TABLE H-2
PERCENT CARBON DISTRIBUTIONS AS DETERMINED BY
PHASE-DISCRIMINATING PROBE
TEST NO. 5
Iso-Octane
= 0.65
Pressure = 1 atm
Inlet Swirl No. = 0.6
Inlet Air Temperature = 533°K
D = 12.23 cm
R = 6.17 cm
o
X/D = 1.21
R/Ro
.671
.617
.252
-.287
-.611+
-.791
X/D = 2.60
.671+
.67^
.251+
.06U
.061+
-.238
-.1+70
-.785
Percent Carbon
THC
.211+
.25!+
1.1+6
.972
.76
.373
.20k
1.127
1.01
.93
.63
.67
VHC
.063
.155
.U55
.097
.566
.35
.225
.21+1
A5
.326
.331
.21+7
.335
.21+7
THC
.225
.239
5.^9
;592
.701+
.376
1.06
.79
l.Ol
.62
VHC. THC
.059
.135 -239
.51
.51
.31
.236
.36
.3M+
.326
.362
.3^+
VHC
.061
.279
.29
.1+62
iZ9
-------
TABLE H-2 (Cont'd)
PERCENT CARBON DISTRIBUTION AS DETERMINED BY
PHASE-DISCRIMINATING PROBE
Iso-Octane
i = 0.65
Pressure = 1 atm
Inlet Swirl No. = 0.6
Inlet Air Temperature = 533°K
D = 12.23 cm
X/D = 3-99
R/RO
.56^
.56H
.k60
.323
.050
-.255
-.5039
-.5039
-.785
--785
X/D = 5.38
.669
.363
.139
.067
-.026
-.212
-.299
-.^59
RQ= 6.17 cm
Percent Carbon
THC
.7^6
.62
.89^
.120
1.17
-^37
.986
.60
.528
.613
.775
1.197
= 915
,986
VHC
.14-21
396
.797
1.12lf
.858
.20U
.209
l.M*
.97
.60
.53
.613
.775
1.20
92
.98
THC
.683
.889
.718
1.13
l.Ul
.63U
.528
.669
1.06
1.13
.986
.986
VHC THC
A29 .65
.i+00
.822
1.02
.923
.225
.196
IM
1.06
.63^
.528 .U93
.67
1.06
i.l
99
.98
VHC
A05
.14-05
.997
.^
.93
130
-------
TABLE H-3
PERCENT CARBON DISTRIBUTIONS AS DETERMINED BY
PHASE-DISCRIMINATING PROBE
TEST NO. 6
No. 2 Fuel Oil
§ = 0.65
Pressure = 1 atm
Inlet Swirl No. = 0.3
Inlet Air Temperature
D = 12.23 cm
R = 6.17 cm
o
533°K
X/D = 1.21
R/Ro
.673
.14-58
.263
.168
.Olio
-.155
-.279
-.366
-.569
-.786
-.786
Percent Carbon
THC
.0809
.36
.67
1.36
1.79
1.-68
.895
.918
.5U2
.317
.5^
VHC
-0396
31
.689
.753
1.68
l.i*8
.60k
.593
-U03
.21k
THC
.122
.307
.808
1.0
1.89
1.81+
.781
.827
A78
.1+21
VHC
.0335
.376
.732
.739
1.5
1.33
A87
.613
.309
.282
THC
.7^8
1.18
1.5^
1.65
.77^
.7^9
.510
VHC
.723
1.60
.3^0
.322
131
-------
TABLE H-3 (Cont'd)
PERCENT CARBON DISTRIBUTIONS AS DETERMINED BY
PHASE-DISCRIMINATING PROBE
No. 2 Fuel Oil Inlet Swirl No. = 0.3
$ = 0.65 Inlet Air Temperature = 533°K
Pressure = 1 atm D = 12.23 cm
RQ = 6.17 cm
X/D = 2.60
R/R0
.671
.671
.1+52
.1*52
.26U
.130
.130
.056
-.014-5
-.0*4-9
-.158
-.158
-.252
.252
-.582
-.766
-.766
Percent Carbon
THC
.288
-385
.185
.20
.313
1.02
1.31
1.59
1.91
1.08
.512
.713
A3U
.569
.753
VHC
.25U
.239
.251*
.781
1.05
1.05
1.1*8
1,1*3
1.72
1.29
1.26
.7
.596
.3^8
.1+05
537
THC
.38
.11*8
.169
.293
.87
1.37
1.69
1.08
512
.591*
.388
.778
.735
VHC
.2U8
.23U
.351
1.12
1.29
1.9
1.25
.658
.582
.322
.521
531*
THC
.39
.191
.91
1.25
1.12
.523
.61
.322
.87
VHC
.272
.198
.U26
1.12
1.32
1.16
.626
.298
.585
132
-------
TABLE H-3 (Cont'd)
PERCENT CARBON DISTRIBUTIONS AS DETERMINED BY
PHASE-DISCRIMINATING PROBE
No. 2 Fuel Oil Inlet Swirl No. = 0.3
$ = °-65 Inlet Air Temperature = 533°K
Pressure = 1 atm D - 12.23 cm
Ro - 6.17 cm
X/D = 3-99
R/Ro
.666
.1+25
.2U6
.21+2
0
-.00k
-.175
-.1*25
-.1+1+8
-.571
-.798
-.798
X/D = 5.38
.676
.1+1+9
.322
.088
.027
.010
.010
-.11+9
-.326
-.51*8
-.788
THC
.3C4
.0713
.1^5
.092
.968
.1*87
.0593
.0531
.169
1.52
1.142
.211+
.0821
.01514-
.1*35
.38
.3^3
.0232
.258
.367
VHC
.0739
.0361
.137
.13
.115
.902
.1*26
-0536
.136
l.ll
1.2U
.133
.0858
.0292
-363
352
.598
.683
.21*9
.0213
.089^
.127
Percent
THC
.118
.11
.1*2
.021*1
.109
1.81
.293
.065
.0118
.3^
.221
.019^
.182
.^02
Carbon
VHC THC
.109 .195
.0^62
.126
.858
.1+1
.156
1.25 1.U7
.11*9 -2^3
. 072 . 061*6
.0253
.296
.666
.211
.0171
.061+ .15!+
.11*1*
VHC
.116
3.11
.162
.572
.051
133
-------
TABLE H-14-
PERCENT CARBON DISTRIBUTIONS AS DETERMINED BY
PHASE-DISCRIMINATING PEOBE
No. 2 Fuel Oil
$ = 0.65
Pressure = 1 atm
X/D = 1.21
Test No. 8
Inlet Swirl No. = 0.3
Inlet Air Temperature = 755°K
D = 12.23 cm
RQ = 6.17 cm
Percent Carbon
R/R
o
.675
.675
.14-61*
.252
.252
.135
.060
-.055
-.158
-.261
-.373
-.373
-.14-60
-.579
-.67k
-.787
-.787
THC
.31
.27
.817
.611*
1.1*8
3.1+3
3.1U
2.53
1.32
.255
.185
.671
.119
.9^6
l.M*
1.U06
VHC
.281
.333
.257
.81*1*
1.1*6
3.07
3.08
2.89
1.105
.158
.356
.133
.503
.668
.656
THC
.781*
1.51
3.07
3.10
2.50
1.09
.187
.1*85
.12
.771
987
1.1*1
VHC
.326
.262
.710
3.1^
2.92
2.71
1.09
.171*
.323
.112
.428
.656
THC
.637
3.05
.206
.169
1.21*
VHC
.323
32
.726
2.72
.166
.669
134
-------
TABLE E~h (Cont'd)
PERCENT CARBON DISTRIBUTIONS AS DETERMINED BY
EHA.SE-DISCRIMINATING PROBE
No. 2 Fuel Oil
$ = 0.65
Pressure = 1 atm
Inlet Swirl No. = 0.3
Inlet Air Temperature = 755 K
D = 12.23 cm
RQ = 6.17 cm
X/D = 2.60
Percent Carbon
R/R0
.661+
.U7l
.1*71
.271
.271
-oi^
-.076
-.082
-.175
-.251
-.251
-.388
-.569
-.783
THC
.312
.061
.098
.137
1.01
1.26
1.15
.896
.Ul
.1^7
.09
.218
.521+
VHC
.166
.lUi
-379
.265
1.01+
1.31
1.25
.68
.kk6
A86
.183
.223
.268
THC
.313
.0916
.133
.98
l.ll
.895
.1+63
.061
.199
.566
VHC THC
.201+
.139 .0787
.375
.258
1.1 1.08
.752 .88
.528 M
.156
.212
.3^
VHC
.226
.158
.218
.838
.58
.38
135
-------
TABLE H-1+ (Cont'd)
PERCENT CARBON DISTRIBUTIONS AS DETERMINED BY
PHASE-DISCRIMINATING PROBE
No. 2 Fuel Oil
$ = 0.65
Pressure = 1 atm
Inlet Swirl No. = 0.3
Inlet Air Temperature = 755° K
D = 12.23 cm
Ro = 6.17 cm
X/D = 3-99
R/Ro
.672
.U66
.211
.211
.028
-.062
-.192
-.218
-.232
-.390
-.390
-.596
-.79^
Percent Carbon
THC
.0121+
.00298
.1U5
.653
.29
.235
.235
.172
.0063
. 00675
. 00^3^
.15U
VHC
.0121+
.0035
.137
-117
.639
.273
.182
.189
.21
. 00681
.001+5
.103
THC
.011+1+
. 00266
.153
.596
. 00636
. 001+17
.15
VHC THC
.0202 .0158
.128 .129
.523 .566
.18U
.00836 .00888
. 001+66
.115
VHC
.0196
.120
136
-------
TABLE E-h (Cont'd)
PERCENT CARBON DISTRIBUTIONS AS DETERMINED BY
EHASE-DISCRBTOIATING PROBE
No. 2 Fuel Oil
= 0.65
Pressure = 1 atm
Inlet Swirl No. = 0.3
Inlet Air Temperature = 733°K
D = 12.23 cm
R = 6.17 cm
o
X/D = 5.38
R/R0
.676
-Mi-9
.2U6
.213
.113
.033
-.028
-.125
-.205
-.389
-.57^
-.786
-.786
THC
. 00075
. 00188
-00573
.0063
.513
.23
.23
358
.102
. 00171
.003*19
.026^
.0186
VHC
. OOU25
.00188
.OlU
-0132
.Uoi
-355
.309
.375
.107
.00276
.001^8
.0381+
Percent Carton
THC VHC THC VHC
.0029^
.0128
. 00883
.273
.278
.065
.0013^ .00251
.002li.3 .00593 '°°5
.0232 .0^70 -0235 .0^8
137
-------
TABLE H-5
PERCENT CARBON DISTRIBUTIONS AS DETERMINED BY
PHASE-DISCRIMINATING PROBE
TEST NO. 11
Propane
$ = 0.65
Pressure =
X/D = 1.21
R/R0
.67k
-^35
.250
.150
.057
-.362
--575
-.783
X/D = 2.60
.676
.676
.635
.580
.^56
.267
. oUi
-.166
'-.360
-.1462
-.577
-.669
; 1 atm
THC
.U29
12.83
36.77
to. 55
35.51
35-19
15.37
.268
.133
.123
.387
2.8l
8.32
12.12
13.7^
15.21
12.0*4
11.13
1.63
0.0
Inlet Swirl No. = 0.3
Inlet Air Temperature = 533°K
D = 12.23 cm
RQ = 6.17 cm
Percent Carbon
VHC
.11*
5.09
26. k3
3,k.k2.
32.89
32.57
13.98
.08
.37*4
.57
1.57
5.08
10.1*4
10.22
12.67
9.35
7.93
1.37
.0169
THC
3.61
9.86
27.3*4
*kL.32
37.36
31.03
12.16
.118
1.02
.729
1.66
8.11
12.67
11. 8*4
15.96
15.01
2.13
-0103
VHC
.0387
5.56
27.17
31.17
33.16
28.7*4
7.22
.273
.280
.68
2.14.9
7.61+
9-27
11.01
12.12
10. k6
7.95
1.39
0.0
THC VHC
.257
28.655 27.3*4
36.55 32.62
12.95
.0868 .118
-131
1.38
7.97
15.37
1.25
138
-------
TABLE H-5 (Cont'd)
PERCENT CARBON DISTRIBUTIONS AS DETERMINED BY
PHASE-DISCRIMINATING PROBE
Propane
§ = 0.65
Pressure = 1 atm
Inlet Swirl No. = 0.3
Inlet Air Temperature = 533°K
D = 12.23 cm
RQ = 6.17 cm
X/D = 3.99
R/R0
.6kl
.55k
.37k
.263
.026
-.oik
-.170
-.365
-.365
-.705
-.705
Percent Carbon
THC
.0168
.19
k.02
2.29
9-93
5.1
7.25
6.67
.kQ6
k39
VHC
,026
.3
1.^3
.075
.075
.098
.098
THC
.0365
3-6U
3.05
10.26
5-25
7.7^
6.kk
.503
.k$l
VHC
.2
.Qk
.075
.0709
THC
.Okk
2.70k
k.9k
7.69
6.76
-^39
VHC
.0281
.0709
139
-------
TABLE H-5 (Cont'd)
PERCENT CARBON DISTRIBUTIONS AT DETERMINED BY
PHASE-DISCRIMINATING PROBE
Propane
§ = 0.65
Pressure = 1 atm
Inlet Swirl No. = 0.3
Inlet Air Temperature = 533°K
D = 12.23 cm
RQ = 6.17 cm
X/D = 5-38
R/R0
.63k
.63k
.63^
571
.Mf5
.kk5
.1+29
.2^0
.033
-.152
-.370
-.U65
-.582
-.68k
-.782
THC
.113
.051+
.0^91
.87
^.58
k.kQ3
6.01
7.98
7.3
if. 08
3.2k
1.17
.105
.0232
VHC
.09
1.01
if. 56
3 Ml
3.39^
5.35
7.59
5.8i
3.53
2.72
1.07
.159
-0387
Percent
THC
.Oif97
.05^9
.68
.kk9
If. 66
6.15
7.68
7-02
3.95
2.80
1.228
.109
.0287
Carbon
VHC
1.0k
3.97
3.799
5-77
6.79
5-77
3.52
1.2U
.156
.0383
THC
-0677
.0^52.
.kk7
5.99
1.67
-0321
VHC
,if62
Ik 02
7.18
.038
140
-------
APPENDIX I
TABULATED SPECIES CONCENTRATION DATA
141
-------
TABLE 1-1. SPECIES CONCENTRATION DISTRIBUTIONS - TEST NO. 1
Iso-Octane
$ = 0.65 ± .01
Pressure = 1.0 atm
R/R0
-0.80
-0.58
-o.i*8
-0.38
-0.17
o.oi*
0.2l*
0.1*5
0.66
0.85
0.88
R/RO
-0.80
-0.58
-0.38
-0.17
o.oi*
0.2l*
0.1*5
0.66
0.85
0.87
Inlet Air Swirl No. = 0.3
Inlet Air Temperature = 530 - 3°K
X/D = 1.21
3
23
1*1
58
75
71
70
1*2
6
1
-
X/D = 1.21
5
36
_
_
71
-
69
23
8
-
NO, ppm
2.60
1
22
-
91
127
112
127
68
il*5
-
3
NOXJ ppm
2.60
13
k6
119
_
_
-
89
35
-
13
1*.00
6
1*1*
_
101
il*6
116
158
101
26
3
-
i*.oo
2k
66
127
165
-
Ill
1*5
li*
_
5.38
_
-
_
-
167
133
172
73
13
3
-
5.38
_
-
_
_
186
ioi*
38
13
_
142
-------
TABLE 1-1. SPECIES CONCENTRATION DISTRIBUTIONS
(Continued)
R/Ro
-0.80
-0.58
-0.1*8
-0.38
-0.17
o.oU
0.2l*
0.1*5
0.66
0.85
0.87
R/R
0
-0.80
-0.58
-0.1*8
-0.38
-0.17
o.oi*
0.2l*
0.1*5
0.66
0.85
0.87
R/RO
-0.80
-0.58
-0.1*8
-0.38
-0.17
o.oi*
0.2k
0.66
0.85
0.87
CO,
X/D - 1.21
0.90
3.00
5.80
8.10
11.1*0
12.20
10.6
6.20
1.80
0.78
-
C02
X/D = 1.21
2.3
6.5
7-6
7.6
7.6
7.2
7.6
1-k
5.0
1.8
"
02,
X/D = 1.21
16.85
9.00
1*.68
2.1*2
0.65
0.52
1.22
14.81
11.96
7.50
-
Mole %__
2.60
0.56
1.18
_
3.50
7^50
10.00
5.95
2.10
0.80
-
0.60
, Mole %
2.60
3A
7.9
-
9.1*
9.0
8.0
9-k
9.0
6.2
-
3-5
Mole io
2.60
15.86
10. ll*
-
3.77
0.78
0.29
1.17
5 = 59
11.57
-
ll*.l*3
k.oo
0.51*
0.85
^
2.10
6.80
10.20
5.30
1.80
0.80
0.50
-
IK oo
5.3
8.1
-
10.2
9.1*
8.0
10.2
10.0
7.0
3*k
k.oo
Ik. CO
8.15
k.3k
-
0.95
0.26
1.03
I*. 1*3
10.20
16.
-
5.38'
_
tm
5.60
10.00
3.90
0.75
o.kQ
0.38
-
5.38
-
-
-
-
10.0
8.6
10.7
9-k
6.2
3.3
5.38
-
-
"
1.15
0.1*0
1.65
6.75
12.13
16.20
143
-------
TABLE 1-2. SPECIES CONCENTRATION DISTRIBUTIONS - TEST NO.
Iso-Octane
0 = 0.66
Pressure =3.3 atm
R/R
' 0
-0.8k
-0.80
-0.66
-0.57
-0.45
-0.28
-0.24
-0.04
0.17
0.38
0.58
0.69
0.76
0.80
R/R
0
-0.81*
-0.80
-0.66
-0.57
-0.1+5
-0.28
-0.24
-0.04
0.17
0.38
0.58
0.69
0.76
0.80
NO,
X/D = 1.21
3
-
58
92
-
68
-
70
63
81
90
71
12
7
NOX,
X/D = 1.21
13
-
74
-
-
_
_
-
-
-
-
-
17
12
Inlet
Inlet
ppm
2.60
-
10
47
-
90
-
74
58
-
91
94
-
-
6
ppm
2.60
-
-
60
-
-
_
_
-
-
_
109
-
-
_
Air Swirl
No. = 0.3
Air Temperature = 533 - 4°K
4.00
2
-
66
-
76
_
55
5^
60
75
71
-
-
6
4.00
10
_
91
_
_
_
_
-
_
_
12
_
-
_
5-38
-
3
18
-
73
-
66
50
53
79
45
-
15
-
5.38
_
10
32
_
_
_
_
_
_
_
70
_
29
_
144
-------
TABLE 1-2. SPECIES CONCENTRATION DISTRIBUTIONS
(Continued)
CO, Mole
B./RO
-0.8H
-0.80
-0.66
-0.57
-O.U5
-0.28
-0.21+
-o.oU
0.17
0.38
0.58
0.69
0.76
0.80
/
R/R0
-0.81+
-0.80
-0.66
-0.57
-0.1+5
-0.28
-0.21+
-O.Qlf
0.17
0.38
0.58
0.69
0.76
0.80
X/D = 1.21
0.00
-
2.10
7.30
-
9.8
-
9.6
10.05
8.70
7.00
3.35
0.36
O.UO
C02,
X/D = 1.21
0.5
H
7.3
7.95
-
6.1
_
5-9
5.9
6.3
8.15
8.1
2.6
3.25
2.60
-
0.12
0.80
-
5.5
-
8.85
9.^
-
7.55
3.15
-
-
0.12
Mole °,
2.60
-
1.95
6.0'
-
8.6
-
6.95
6.7
-
7.75
8.85
«.
1.6
i+.oo
0.25
-
3. to
7.30
-
-
8.5
8.6
7.75
6.9
1+.2
-
-
0.30
i
14.00
2.1
-
8.8
-
8.0
-
7.1+2
7.5
8.0
8.^5
9.^
-
2.1+5
5.38
_
0.05
o.i+o
-
1+.6
-
7A
8.U
8.1
5^92
1.62
-
.3.50
-
5-38
-
0.85
U.i
-
10.2
-
8.1+
7.25
7-75
9.3
8.75
3.65
145
-------
TABLE 1-2. SPECIES CONCENTRATION DISTRIBUTIONS
(Continued)
02, Mole
n/no
-0.8^
-0.80
-0.66
-0.57
-0.1+5
-0.28
0.2k
o.ok
0.17
0.38
0.58
0.69
0.76
0.80
X/D = 1.21
20.38
-
7.83
0.37
-
0.07
0
0
0.13
0.75
U.72
16.6
17.27
2.60
_
18.39
ll.8o
-
0.87
-
0.25
0.12
-
0.37
2.98
_
_
19.1?+
h.oo
18.23
-
3.00 (
-
0.18
-
0.26
0.13
0.13
0.13
1.78
_
17.69
5.38
_
19.61
15.05
-
0.63
-
0
0.25
0.06
0.18
6.37
M
15,56
_
146'
-------
TABLE 1-3. SPECIES CONCENTRATION DISTRIBUTIONS - TEST WO.
Iso-Octane
0 = 0.65
Pressure = 1.0 atm
R/R0
-0.85
-0.66
-0.1+5
-0.2k
-o.dk
0.17
0.38
0.1+8
0.58
0.80
R/R0
-0.85
-0.66
-0.1+5
-0.21+
-o.ok
0.17
0.38
0.^8
0.58
0.80
__ /_
R/RO
-0.85
-0.66
-0.1+5
-0.21+
-o.dk
0.17
0.38
0.1+8
0.58
0.80
Inlet Air Swirl
Wo. = 0.6
Inlet Air Temperature = 530 ~ 5°K
NO
X/D = 1.21
3
1+2
87
71
65
63
-
90
50
1
NO
X/D - 1.21
18
72
-
-
-
-
-
_
-
18
CO,
X/D = 1.21
0.3
2.0
9.55
ll.l
11.25
11.25
-
8.15
k.7
0.35
> PPm
2.60 i+.oo
!+ k
35 12
27 23
22 103
16 76
19 81
25 160
-
86 55
5 3
x-, PPm
2.60 i+.oo
10 10
1+1+ 22
_
-
-
-
-
195
87 65
7 ll+
Mole %
2.60 i+.oo
0.1+5 O.U2
i.oo 0.50
6.6 5-2
9.7 12.6
11.25 13-5
11.1+ 13.2
9.0 8.9
2.1
1.1+5 k.O
0.1+3 o.l+
5.38
8
39
187
99
72
80
127
-
79
k
5.38
ik
1+1+
-
-
-
-
-
83
6
5-38
o.i+o
0.25
5-7
12.5
11.7
13.2
ll.l
0.15
1 i
O.U5
147
-------
TABLE 1-3. SPECIES CONCENTRATION DISTRIBUTIONS
(Continued)
R/R,
C02, Mole
' U
0.85
0.66
0.1+5
0.21+
o.oi+
0.17
0.38
O.U8
0.58
0.80
R/R
0
0.85
0.66
0.1+5
0.21+
O.oi+
0.17
0.38
O.U8
0.58
0.80
X/D = 1.21
1.35
8.0
7.5
6.7
6.1+
6.55
-
8.1
8.3
1.9
02,
X/D = 1.21
18.81+
5.^7
0.05
o.oo
0.00
0.00
-
0.38
3.31
18.33
2.60
2.6
7-^5
7.9
7.0
6.05
6.6
7-6
-
8.5
3.5
Mole %
2.60
17.33
8.78
1.27
1.01
1.11+
1.39
1.27
-
5.95
16.02
i+.oo
2.1+5
5.^5
9.8
7.3
6.8
6.85
8.65
10.75
7.0
2.6
i+.oo
17-1+6
13.69
i.il+
0.05
0.02
0.02
0.25
2.5
9.0
17.^3
5.38
3.5
6.6
9-95
7.05
6.3
6.7
7.5
-
9.5
3.8
5.38
16.26
11.59
0.51
0.18
2.1+0
0.08
0.63
_
9.33
15.75
148
-------
TABLE I-14-- SPECIES CONCENTRATION DISTRIBUTIONS - TEST NO. 8
No. 2 Fuel Oil
0 =0.67
Pressure = 1.0 atm
Inlet Air Swirl No. = 0.3
Inlet Air Temperature = 755 K
NO, ppm
R/Ro
-0.81+
-0.80
-0.66
-0.57
-0.14-5
-0.38
-0.214.
-0.17
-O.Ol*
o.oi+
0.17
0.25
0.38
0.1+5
0.58
0.66
0.80
0.81+
R/Ro
-0.81+
-0.80
-0.66
-0.57
-0.1+5
-0.38
-o.2k
-0.17
-0.01+
o.oi+
0.17
X/D = 1.21
ll+
-
31
-
1+7
-
79
-
108
_
108
-
56
-
l+l
-
11+
-
X/D = 1.21
19
-
38
-
61).
-
-
-
-
56
2.60
-
23
-
1+9
-
83
-
15!+
-
188
-
188
-
135
-
83
-
55
NOX, ppm
2.bO
-
32
-
57
95
-
U.oo 5.38
103
39
157
101
209
114-3
295
196
.290
201
280
203
225
151
-
119 157
-
8U
J+.oo 5'38
-
73
- -
105
114-6
"
*
149
-------
TABLE I -1+. SPECIES CONCENTRATION DISTRIBUTIONS
(Continued)
NOX, ppm
' w
0.25
0.38
0.1+5
0.58
0.66
0.80
0.81+
R/R0
0.81+
0.80
0.67
0.58
0.1*5
0.38
0.21+
0.17
o.oU
0.01+
0.17
0.25
0.38
0.1+5
0.58
0.66
0.80
0.81+
X/D = 1.21
_
83
-
-
-
27
-
X/D = 1.21
1.1+7
-
1.11
-
1.9
-
3.99
-
6.19
_
6.05
-
2.82
_
1.52
-
0.97
-
2.60
_
-
135
-
83
-
-
CO, Mole °1(
2.60
-
0.86
-
0.87
-
1.21
-
3.0
-
7.18
-
5.18
-
1.65
_
0.83
-
0.63
i+.oo
_
-
-
-
119
-
-
r
j
l+.OO
-
0.77
-
0.62
-
1.06
-
3.1+3
-
8.1+6
-
!+.l+3
-
1.1+6
-
0.77
-
0.55
5.38
_
225
-
-
157
-
-
.5.38
0.38
-
0.31+
-
0.8
-
2.6
-
5.88
_
1+.06
_
1.25
_
0.65
_
_
150
-------
TABLE I-It-. SPECIES CONCENTRATION DISTRIBUTIONS
(Continued)
C02, Mole
/ 0
-0.81+
-0.80
-0.67
-0.58
-0.1+5
-0.38
-0.21+
-0.17
-o.oU
o.oi+
0.17
0.25
0.38
0.1+5
0.58
0.66
0.80
n.RU
X/D = 1.21
1+.01
-
5.07
-
6.07
-
7.81
-
7 = 92
_
7.98
-
7-1*
-
5.79
-
1+.7
fm
2.60
_
6. ill-
-
6.95
-
7.85
-
8.25
-
8.73
-
9.3
-
8.M+
-
7.1
-
6.1
i*.oo
_
7.39
-
8.21+
-
8.96
-
9.7
-
8.65
-
9.25
-
8.87
-
8.1+9
-
7.1+1
5.38
7.^3
-
8.3
-
9.0
-
9-7^
-
9.28
-
9-53
-
9.3
-
8.56
-
151
-------
SPECIES CONCENTRATION DISTRIBUTIONS
(Continued)
R/R0
0.81+
0.80
0.67
0.58
0.1+5
0.38
0.21+
0.17
0.01+
0.01+
0.17
0.25
0.38
0.1+5
0.58
0.66
0.80
0.81+
°2>
X/D - 1.21
11+.7
-
13.3
-
10.8
-
5.3
-
2.9
-
3.2
-
7.U
-
12.0
-
11+.2
-
Mole %
2.60
_
12.1+
-
10.8
-
8.7
-
3.7
-
1.1
-
1.7
-
6.9
_
10.5
_
12.1+
4.00
_
9.8
-
8.15
-
6.1
-
1.7
-
0.38
-
2.1
-
5-9
_
7.5
_
9-7
5.38
10.1
-
8.0
-
6.2
_
2.5
-
1.1
-
2.0
-
5.1
-
H
7.1*
_
_
152
-------
TABLE 1-5- SPECIES CONCENTRATION DISTRIBUTIONS - TEST NO.
No. 2 Oil
0 - 0.67
Pressure =1.0 atm
R/RO
-0.81+
-0.80
-0.66
-0.58
-0.1+5
-0.38
-0.2k
-0.17
-o.dk
o.ok
0.17
0.25
0.38
0.58
0.66
0.80
o 81+
Inlet Air Swirl No. = 0.3
Inlet Air Temperature = 529 - 6 K
X/D = 1.21
_
_
_
-
-
-
-
-
87
_
83
_
1+6
51
21
30
5
15
NO, ppm
2.60
-
19
-
29
-
52
-
97
-
112
101
93
-
1+2
-
21
-
17
U.oo
-
37
-
66
-
95
-
120
-
103
-
105
-
86
-
53
-
1+5
5-38
39
-
62
-
119
-
132
-
132
-
119
110
102
*
65
«
153
-------
TABLE 1-5. SPECIES CONCENTRATION" DISTRIBUTIONS
(Continued)
X/D = 1.21 2.60 l+.OO 5.38
-O.Qk
-0.80
-0.66
-0.58
-0.1+5
-0.38
-0.21+
-0.17
-o.oi+
o.ok
0.17
0.25
0.38
0) i CT
4?
0.58
0.66
0.80
0.81+
R/R0
-0.8^
-0.80
-0.66
-0.58
-0.1+5
-0.38
-0.2k
-0.17
-0.01+
o.tik
0.17
0.25
0.38
0.1+5
0.58
0.66
0.80
0.81+
-
_
-
6k
-
-
-
-
-
-
-
66
38
-
19
-
CO,
X/D = 1.21
L.k6
1.3
-
2.82
_
5.07
-
6.57
_
5.82
_
2.97
_
1.52
_
0.99
_
32
_
36
-
58
-
-
-
-
-
-
lift
^+0
26
-
23
Mole %
2.60
_
1.28
_
1.17
-
1.1+9
_
3.^7
_
7-1+8
k.3
2.8
-
1.09
-
0.89
-
1.13
l+l
-
69
-
-
-
-
-
-
-
-
-
-
-
-
-
i+.oo
fm
1.08
_
0.97
_
1.23
_
3.79
_
5.83
-
3-7
_
1.51
_
1.10
_
1.22
:
66
-
-
-
-
-
-
-
-
-
-
-
-
-
5.38
0.62
_
0.66
_
1.M+
«
2.6
n
^.5
3.36
1.39
1.03
H
0.55
154
-------
TABLE 1-5. SPECIES CONCENTRATION DISTRIBUTIONS
(Continued)
R/RO
0.81+
-0.80
0.66
0.58
-0.1+5
-0.38
-0.2k
-0.17
-o.oi+
o.oi+
0.17
0.25
0.38
0.1+5
0.58
0.66
0.80
0.81+
R/R0
-0.81+
-0.80
-0.66
-0.58
-oA5
-0.38
-0.2k
-0.17
-o.ok
o.ok
0.17
0.25
0.38
oA5
0.58
0.66
0.80
0.81+
c
X/D - 1.21
3.01
-
5.11
M
6.91+
-
8.35
-
8.1+5
-
8M
-
7.33
-
5 A3
-
3.17
-
X/D = 1.21
15-9
_
13.1
-
8.5
-
3.2
-
1.7
-
2.9
-
7.8
-
12.6
_
16 A
-
02, Mole °,
2.60
_
5A7
_
6.1k
-
7A8
-
9.12
-
8.87
8.86
8.96
-
7.15
-
5.65
-
5-31
02, Mole
2.60
-
12.6
-
11.9
-
9.1
-
k.o
-
0.89
-
-
-
10.1
-
13.0
-
13 A
6
4.00
6.53
7.73
_
8.92
-
9.62
_
8.87
-
9.39
-
8.87
-
7,98
-
6.95
%
k.oo
-
11.6
-
9.2
-
6.3
-
2.2
-
2.5
-
3.58
-
5.9
-
8A
-
10.1+
5.38
6.6
^
8.06
9.36
_
9.65
_
9A5
_
9.2l+
-
9.05
_
8.73
-
7.28
-
5.38
11.7
-
8.8
-
k.9
-
3.3
-
1.9
-
3.3
-
5-7
-
6.6
-
10.6
-
155
-------
TABLE 1-6. SPECIES CONCENTRATIONS DISTRIBUTIONS - TEST NO. 11
Propane
d = o.6k i .01
Pressure = 1.0 a tin
R/R
/ o
-0.80
-0.58
-0.38
-0.17
o.ok
0.2k
O.U5
0.59
0.66
0.70
0.85
R/R0
-0.80
^ -0.58
-0.38
-0.17
0.0**
0.2k
0.59
0.66
0.70
0.85
Inlet Air
Inlet Air
Swirl No. = 0.3
Temperature = 521 - 8° K
NO, ppm
X/D = 1.21
5
8
k7
55
5k
k8
23
-
2
-
1
X/D = Iv21
-
11
-
-
_
-
p.
3
-
k
2.60
3
20
70
70
71
69
68
-
7
-
3
NOX, ppm
2.60
_
-
-
_
_
-
20
-
5
k.oo
ill
75
75
7k
68
70
-
72
-
10
2
k.oo
23
-
-
_
-
-
1U
1+
5.38
k
16
8k
6k
61
78
78
-
10
-
3
5.38
1^
_
31
_
mm
-
~_
Ik
_
7
156
-------
Table I-6. SPECIES CONCENTRATIONS DISTRIBUTIONS
(Continued)
R/RO
-0.80
-0.58
-0.38
-0.1?
O.o4
0.24
0.45
0:59
0.66
0-70
0.85
R/R0
-0.80
-0.58
-0.38
-0.17
0.04
0.24
0.45
0.59
0.66
0.70
0.85
R/R0
-0.80
-0.58
-0.38
-0.17
o.o4
0.24
0.45
0.59
0.66
0.70
0.85
X/D = 1.21
0.1
2.4
7-4
7.3
7.3
7.4
6.25
-
0.55
-
0.0
X/D = 1.21
0.9
5.05
6.15
5.55
5.55
5.6
5.8
-
1.8
-
0.15
0
X/D = 1.21
19.71
10.94
1.00
0.2
0.37
1.12
2.98
-
17.99
20.73
CO, Mole
2.60
_
2.36
-
9.3
9.4
9.08
8.58
-
2.8
_
0.22
COo, Mole
C- y
2.60
0.4
4.84
-
5.92
5.92
5.99
6.47
-
5.51
-
0.93
, Mole
-------
TABLE 1-7. EXHAUST SPECIES CONCENTRATIONS (iSO-OCTAKE)
Test Swirl
No. No.
l 0.3
0.3
0.3
0.3
0.3
0.3
2 0.3
0.3
0.3
0.3
0.3
0.3
0.3
0.3
0.3
3 0.3
I
k 0.3
5 0.6
0.6
0.6
Pressure
(atm)
1.0
1.0
1.0
1.0
1.0
1.0
1.0
1.0
1.0
1.0
1.0
1.0
1.0
1.0
1.0
3.5
3.5
1.0
1.0
1.0
Inlet Air
Temperature
7^0
7^6
yli ^
yji li
7^6
765
7^9
7^7
637
638
530
528
527
533
52U
527
532
530
535
525
0.89
0.65
0.57
0.73
0.8l
0.96
0.73
0.65
0.65
0.72
0.57
0.61
0.72
0.80
0.88
0.65
0.65
0.65
0.62
0.67
Apinj
( atm)
\ /
"
"
~
"~
7.71
**
"
7.65
"
7.6k
9.99
12.72
-
~
158
-------
TABLE 1-7. EXHAUST SPECIES CONCENTRATIONS (ISO-OCTANE)
(Continued)
Test
No.
1
2
3
^J
1+
5
02
(Mole $)
U.2
9.6
11.8
7.6
6.0
3.3
6.6
9.*+
9.U
X
7-5
12.1
9.5
7.6
5.9
ij.,6
10.8
10.2
10.9
11.6
10. U
C02
(Mole %)
9-6
7.3
5.9
8.1
8.8
9-9
9.0
8.0
8.0
9.0
6.1+
7.8
8.8
9.6
10.0
7.1
7.5
6.7
6.1
6.9
CO
(Mole f0)
0.98
0.12
0.11
0.25
0.62
1.30
0.30
0.10
0.15
0.21
0.18
0.17
0.38
0.80
1.35
0'.37
0.25
0.3
0.7
0.2
NO
ppm
218
156
101
183
200
230
185
ll+7
91
125
20
50
72
90
102
31
39
1+6
27
55
NOX
ppm
235
169
110
198
210
250
207
185
117
ll+7
1+8
76
93
113
117
1+1+
i+7
52
Uo
s\.
64
159
-------
TABLE 1-8. EXHAUST SPECIES CONCENTRATIONS (NO. 2 OIL)
Test
No.
1
2
3
k
5
6
Test
No.
1
2
3
1+
6
Swirl
Wo.
0.6
0.6
0.6
0.3
0.3
0.3
0.3
0.3
0.3
0.3
0.3
0.3
0.3
0.3
0.3
0.3
°2
(Mole $)
7.7
8.1
7.2
9.6
7.5
6.1
k.9
3.9
7.7
6.8
8.8
5.2
6.6
8.9
5-8
8.8
8.8
5.7
Pressure
(atm)
1.0
1.0
1.0
1.0
1.0
1.0
1.0
1.0
1.0
3-5
1.0
1.0
1.0
co2
(Mole %
8.9
8.6
9.0
7-9
8.8
9.1
9.k
9.5
9-0
10.1+
9.3
11.2
9.9
9.2
10.7
-
-
-
Inlet Air
Temperature
(°K)
530
537
532
522
532
526
530
525
71+8
533
533
533
533
61+1+
6kk
6kk
CO
) (Mole $)
0.007
0.006
0.01
0.001
0.18
0.52
1.11
1.8l
0.0k
0.06
0.0k
0.10
0.11+
0.06
0.^6
0.08
0.02
0.22
1
0.65
0.63
0.67
0.57
0.61+
0.72
0.79
0.87
o.6k
0.65
0.57
0.7k
0.65
0.57
0.7k
0.65
0.57
0.7!+
NO
ppm
66
62
78
52
8k
103
115
125
166
116
10k
ill
108
80
119
120
9k
152
APinj
(atm)
-
~
"°
~
8 r-?o
73
10.81+
13.22
7.09
"
-
NO
ppm
72
66
82
58
85
105
115
_
167
116
_
109
_
120
9k
152
160
-------
Table 1-9- EXHAUST SPECIES CONCENTRATIONS (PROPANE)
Inlet Air
Test Swirl Pressure Temperature
No. No. (atm) (°K) $
1 0.6
0.6
0.6
2 0.3
0.3
0.3
0.3
0.3
3 0.3
14- 0.3
5 0.3
6 0.3
0.3
0.3
^^ -_J
0.3
^ * ~-J
0 ^
w 3
0.3
Test Op
No. (Mole
-------
APPENDIX J
TABULATED VELOCITY DATA
162
-------
TABLE J-l. AXIAL VELOCITY DATA (M/SEC) - TEST NO.
Iso-Octane/Air Inlet Air Swirl =0.3
0 = 0.65 Inlet Air Temperature = 533 °K
Pressure = 1 atm
X/D = 0.123
R/R0 u U'EMS
-0.86 32.7 13-7
-0.86 34.1 14.9
-0.68 12.39 15-4
-0.68 21.5 19-8
-0.50 -3-1 7-6
0.00 7-4 15-8
0.31 25.6 6.2
0.1*7 -0.8 7-4
0.80 23-9 15.3
0.80 24.9 15.7
0.88 14.8 12.2
0.88 16-5 1^^
0.80 21.6 14.2
0.80 33-5 15-1
0.72 19-9 13-5
0.72 33.2 16.6
0.58 6.4 8.0
0.58 8-9 U~6
0.47 *-2 ^°
0.39 2.4 40
0.31
8.6 5.1
29-3 3.9
0.15
-0.88 18-1 ^
-0.88 29-3 14'9
-0 75 W-2 17
u n oo = 20.2
-O 75 JJ'?
;- f, *. 12.3
-0.63 b'3 17 3
-0.50 -2-1 fi'?
-0-40 -3.99 b'7
-0.28 ll1-8 g,
-0.18 26.1
-0.04 13-3 '\
0.06 23.7 50
163
-------
TABLE J-l. AXIAL VELOCITY DATA (M/SEC) - TEST NO. 1
(continued)
X/D = 0.336
R/R0 U U'
-0.88 22.1 16.6
-0.88 l6.k 13-.0
-0.88 21.7 16.9
-0.88 ill.3 17.5
-0.71 9-0 9-3
-0.71 8.7 , 7-k
-0.71 15.8 17.7
-0.55 9-5 7-8
-0.55 9-6 6.9
-0.55 9-9 6.7
-0.38 12.8 12.6
-0.38 17.7 8.2
-0.38 19.1 8.2
-0.18 13.9 10.9
-0.18 17.7 10.3
0.02 7.3 13.0
0.23 18.8 9.5
0.23 22.6 7.3
O.U3 19.8 5.5
0.^3 18.2 6.2
0.61+ 12.5 5.1+
0.6U 13.5 U-5
0.6k 12.9 5.6
0.76 13.8 5.9
0.76 13.7 7.3
0.76 16.3 10.8
0.93 20.9 11.0
0.93 2U.7 ll.O
164
-------
TABLE J-l. AXIAL VELOCITY DATA (M/SEC) - TEST NO. 1
(continued)
X/D = 0..52U
R/Ro U
-0.81 18.8 5.2
-0.81 18.8 5.1
-0.59 17.1 6.1
-0.59 17-3 6.8
-0.38 ik.h 9.9
-0.18 13-1 13.1
0.00 8.7 10.1
0.23 12.8 11.6
0.^3 16.9 8.U
0.6^ 17.1 5-0
0.8k 18.9 5-A
165
-------
TABLE J-2. AXIAL VELOCITY DATA (M/SEC) - TEST WO. 5
Iso-Octane/Air Inlet Air Swirl =0.6
0 = 0.6 Inlet Air Temperature = 533°K
Pressure = 1 atm
X/D = 0.123
R/RO u u-RMS
-0.88 63.20 16.3
-0.88 67.30 22.9
-0.71 36.90 22.6
-0.55 o.oo 12.5
-0.38 31.80 5.1*
-0.38 32.80 it.8
-0.18 31.10 3.1+
0.02 30.10 3.5
0.23 2U.50 7.8
0.1+3 13.50 21+.0
0.61+ 57.60 13.8
0.6^ 65.70 ih.k
o.8i+ 1+3.20 18.3
-0.79 66.70 16.3
-0.63 35.90 21.8
-0.1*9 1.30 12.1
-0.28 31.1+0 l+.O
0.08 29.50 3.1+
0.13 28.50 3.1
0.33 20.70 10.1
0.33 2.98 8.1
0.51* 33-50 25.6
0-7^ 61.50 il+.8
-0.38 30.60 5.1
166
-------
TABLE J-2. AXIAL VELOCITY DATA (M/SEC) - TEST NO. 5
(continued)
X/D = 0.3_36
u1
RMS
-0.59 29-3 5.6
-0.38 27-7 ^-1
-0.18 2h.b 5-3
0.00 22.6 7-7
0.23 26.4 ^-^
O.U3 22.5 5-0
0.72 7-2 10-7
0.614- 12-0 8.7
0.^7 22.3 5-^
0.31 25-0 ^-2
0.15 22.3 6-2
0.00 17 = 9 6-1
-0.18 23.3 6.0
-0.3^ ' 26'3 h'6
-0.51 25-6 6-^
-0.67 22.6 9-9
167
-------
TABLE J-2. AXIAL VELOCITY DATA (M/SEC) - TEST WO. 5
(continued)
X/D = 0.52k
0.68 16.3 5-^
0.56 17-2 5-6
0-39 16.3 6.6
0.23 15-7 5-9
0.23 11.7 8.8
0.06 17^9 ^-8
0.06 17-2 6.7
-O.io 16.2 U.2
-0.10 17-7 6.7
-0.26 lU.O 5-9
-0.26 17-7 6.0
-0.26 . 15.1
-O.U3 18.8 5-8
-0.^3 20.2
-0=59 18.9 5-9
-0.59 21.7
-0.75 19-2 5-5
-0.75 21.6 3.7
0.23 11-9 8.8
168
-------
TABLE J-2. AXIAL VELOCITY DATA (M/SEC) - TEST NO. 5
(continued)
X/D = 1.64
R/RO u
-0.88 1+4.7 7-9
-0.88 47.3 9-1
-0.69 37.8 7.5
-0.51 26.4 7-3
-0.30 29.1 11.4
-o.io 44.0 12.8
0.11 37-3 12.3
0.33 21.4 7-4
0.52 26.0 7-6
0.72 50.2 6.9
0.72 46.8 10.3
0.84 52.8 7-3
0.84 53-7 8.7
0.64 33.5 9-7
0.02 47.1 12.4
169
-------
TABLE J-3. AXIAL VELOCITY DATA (M/SEC) - TEST NO. 6
No. 2 Fuel Oil/Air Inlet Air Swirl =0.3
0 = 0.65 Inlet Air Temperature = 533°K
Pressure = 1 atm
X/D = 0.123
R/RO u u«
RMS
8770 35-7 13-9
.6926 12.3 12.0
.5082 _2.5 5.6
-3033 9.3 9-3
.0984 19.5 5.9
.1066 21.7 2.8
.3115 -3.4 4,9
.5164 1.2 6.9
0 20.0 7.6
.2295 -0.9 IK 7
.1270 4.9 6.3
.0656 16.0 5.7
8770 20.4 12.7
.8770 32.2 14.5
.6926 11.3 6.7
.6926 12.3
.1803 22.9 3.2
-0779 21.5 4.0
.0246 23.5 3.4
170
-------
TABLE J-3. AXIAL VELOCITY DATA (M/SEC) - TEST NO. 6
(continued)
X/D = 0.336
R/R0 U
U
RMS
.8361 111.7 12.5
.8361 Ul.3 12.2
.63H 19.7 6.8
.6311 19.9 7.1
.4262 20.8 3.5
.2213 20.0 2.6
o 20.6 6.1
.2008 20.0 IKO
.393^ 1^.8 6.8
.5984 4.3 5.6
.8033 22.U 13.8
.8033 ll.l 11-9
.7951 11.2 7.8
.7951 39.^ 9-5
.5902 21.4 IK 3
.3852 21.4
.02^6 22.7
26.U
15.7
.6373 6.7 5.5
171
-------
TABLE J-3. AXIAL VELOCITY .DATA (M/SEC) - TEST WO. 6
(continued)
nm
X/D = 0.336
U'RMS
.8770 18.8 10.2
.8770 27.0 i8.it-
.7131 22.3 5.7
.5)4-92 22.9 ^.6
.3852 22.7 5.9
.2213 1^-5 8.3
.1885 17.8 3.9
-393U 15.8 6.2
.1803 20.8 5.0
0 8.2 8.7
.1885 18.7 5.9
.393^ 16.5 5.2
.598^ 8.2 6.2
.598U iU.6 13.1
.1803 20.7 6.3
.5902 3.1 6.3
-5902 IKS 6.14.
.3852 15.3 7.8
.8770 iU.7 9.3
.5902 6.2 6.5
172
-------
TABLE J-3. AXIAL VELOCITY DATA (M/SEC) - TEST NO. 6
(continued)
X/D = 0.52U
R/RO u U'RMS
.9180 1^.6 ^-8
.9180 25.1 10.2
.7131 16-1 3'8
.5082 16.5 ^-2
.2623 12-8 6-9
.3115 17^ 8-f
.0082 8.2 7.^
-.2295 10.1 5.8
-.6393 13-3
I2-8 3'2
-.3525
173
-------
TABLE J-3. AXIAL VELOCITY DATA (M/SEC) - TEST NO. 6
(continued).
X = 51U.5 nm
X/D = 0.52U
R/R° 5 U'RMS
.8770 18.8 U.7
.8770 21.3 10.2
.6721 16.1+ 3-9
.'4918 18.0 it.it
.3320 22.0 U.l
.1803 .11.3 ' 7.6
-.016)4 ll.l 6.3
-.2295 12.7 6.1
-.Us1^ 16.9 U.9
-.598^ 18.5 3.5
-.7623 19.5 if.8
174
-------
TABLE J-U. TANGENTIAL VELOCITY DATA (M/SEC) - TEST NO. 1
Iso-Octane/Air Inlet Air Swirl =0.3
0 = 0.65 Inlet Air Temperature = 533°K
Pressure = 1 atm
X/D = 0.123
R/Ro * ^'RMS
-0.88 11.20 5.k
-0.88 22.20 8.1
-0.71 8.80 6.0
-0.71 12.00 9-2
-0.51 ^.20 3-9
-0.51 ^-20 3-8
-0.39 S-^O 3-3
-0.18 1-99 2-6
0.13 0.10 2'5
0.23 °-3° 2'5
O.U3 -3-50 2-8
O.U3 -3.80 2.8
0.6k -7.00" ^.8
0.611- ' -11-30 8.3
0.80 -13.50 6.5
0.80 -20.00 9-7
0.23 °'02 3'3
175
-------
TABLE J-U. TANGENTIAL VELOCITY DATA (M/SEC) - TEST NO. 1
(continued)
X/D = 0.336
R/RO ' w ^'RMB
-0.90 13.70 8.3
-0.90 22.60 8.3
-0.90 23.00 9-7
-0.75 8.10 6.3
-0.75 8.10 6.8
-0.75 1*4-.50 11.6
-0.59 U.50 U.l
-0.59 U-.80 U.2
-0.38 3-10 3.0
-0.38 3-00 3.0
-0.18 3.00 3-9
-0.18 2.90 3.3
o.oo -U.io 6.1
0.23 -IK 02 U.i
O.U3 -1.U8 3.3
0.6U -3.60 3.7
0.6U -3.00 3.2
0.76 -U.60 3.6
0.76 -6.10 5.9
0.76 -3.90 3-5
0.88 -7.20 5.1
-0.88 11.Uo 7.7
-0.88 13.50 9.8
-0.90 13.80 9.5
176
-------
TABLE J-5. TANGENTIAL VELOCITY DATA (M/SEC) - TEST NO. 5
Iso-Octane/Air Inlet Air Swirl =0.6
0 = 0.6 Inlet Air Temperature - 533°K
Pressure = 1 atm
X/D =0.123
R/R0 W to'pus
-0.88 6.6 15.0
-0.88 1*6.0 lU.2
-0.11 39.6 1^.8
-0.51 25.2 9.2
-0.30 -l.l 3.2
-0.10 -2.3 ^-2
0.11 -5.7 2.5
0.31 -28.0 7,3
0.52 -35.0 11.7
0.72 -U9.8 10.5
0.1+3 -30.2 7.8
0.23 -10-7 3.9
177
-------
TABLE J-5. TANGENTIAL VELOCITY DATA (M/SEC) - TEST NO. 5
(continued)
X/D = 0.336
w '
0.82 33.2 9-0
o.6k -19.0 6.9
0.6U -28.6 18.5
0.80 -29.6 7.8
o.oo -0.7 3.0
o.oo -0.9 5.2
0.67 7.7 8.2
0.67 21.2 15.1
O.U3 -6.3 2.8
0.23 -1.5 2.8
0.23 -1.3 2.7
178
-------
TABLE J-6. TANGENTIAL VELOCITY DATA (M/SEC) - TEST NO. 6
No. 2 Fuel Oil/Air Inlet Air Swirl =0.3
0 = 0.65 Inlet Air Temperature = 533°K
Pressure = 1 atm
X/D = 0.123
R/RO w fi)lRMS
.8361 13.8 k.9
.8361 13.h 5.6
.6311 13.2 7.8
.6311 22.3 9.5
.^62 1+.3 U.I
.U262 7.1 5.5
.2213 2.9 3.0
.2213 3.5 3.1
.01.6k 3.2 1.8
-.1885 1.8 2.1
-.393^ -1.7 2-5
-.6393 -12.0 7.2
-.81A3 -21.5 6-6
179
-------
TABLE J-6. TANGENTIAL VELOCITY DATA (M/SEC) - TEST NO. 6
(continued)
X/D = 0.336
.8115 5.9 6A
.8115 19.8 5.6
.6311 3-1 3.1
.6311 l^.U 8.1
.U262 1.6 2.2
.2213 1-6 2.1
0. 0. 2.2
-.2295 -.7 2.0
-.k3kh -.k 1.9
-.6393 -1.7 2.0
-.8M+3 -l*. 5 3.9
-.8^3 -16.8 5.5
-.6393 -^.5 6.1
180
-------
TABLE J-7- RADIAL VELOCITY DATA
0 = 0.65
Pressure
= 1 attn
Inlet Air Temperature = 533°K
X/D = 0.336
Fuel Swirl No. Distance From &( ins)
Iso-octane 0 = .6 +.2
I
-.1
-A
-.7
+ = 5
+ .7
, +1.1
0 3 +-8
1
No. 2 Fuel 0
+ .2
-.8
-.8
3 +0-3
+0.9
V
4.0
+1.9
+1+.5
+10.9
-7.1
-13-3
-16.1
-10.0
- 2.1
+8.1
-6.5
-12A
-0.5
+ 7.0
Note: Mean flow is radially outward from centerline.
181
-------
APPENDIX K
DETERMINATION OF SPRAY CONE ANGLE FROM HOLOGRAPHIC DATA
The method used to derive the spray cone angle from the spray surface
coordinates involves a "best fit" of the measured data to the surface coordinates
of right circular cones.
The coordinate system used in the analysis of the data is shown in Fig. K-la.
The center line of the conical surface is parallel to the z axis and the vertix of
the cone lies at an arbitrary point. In general a slight deviation of angle 3
existed between the viewing direction, y' and the y coordinate axis. This devia-
tion produced a slight rotation of the transformed coordinate system (in which
the centerline is parallel to the z axis) and the raw data coordinate system
(Fig. K-lb)). The first step in the data analysis was to assume a value of the
angle of rotation, 3, so that the measured data were transformed from the (x, y1,
z) system to the (x, y, z) system:
y =y'cos/3-z1 sin/3 (K-i)
and
z=y'sin/3 + z'cos/3
(K-2)
Such a transformation was made for several assumed values of (B. The angle, 3,
was varied over an appropriate range as determined from estimates obtained from
raw data plots.
The data in the transformed coordinate system (x, y, z) was assumed to lie
on the surface of the cone whose apex was located at the point (XQ, y , z ) and
whose vertex angle was 9 (see Fig. K-la). The unknowns (x , y , ZQ, 9, 3) were
determined using a technique in which three parameters (x , y , 3) were varied
parametrically and the remaining variables (zo, 8) were obtained by means of
a multiple, linear regression analysis. The set of unknowns yielding the highest
value of the computed coefficient of determination was considered to be the
best fit of the data. Details of this procedure follow.
Any cross-section perpendicular to the z-axis is a circle whose radius is
given by
(K-3)
182
-------
FIG. K-1
COORDINATE SYSTEMS EMPLOYED FOR SPRAY CONE ANGLE DATA ANALYSIS
a) CONE ORIENTATION IN ROTATED COORDINATE SYSTEM
b) ROTATION OF COORDINATE SYSTEM
76-09-2-8
133
-------
This radius is related to the axial coordinate and vertex angle by:
(K-4)
Combining these two equations, squaring both sides of the result, and rearranging
yields the working equation:
zc= z0+ a(x-x0)2+ (y-y0) (K_5)
where the variable, zc, indicates that this equation is used to estimate the
measured value of z for each data point from the measured values (x, y) and the
unknowns (x , y , z , 6, B). The variable, a, is introduced for convenience and
is defined by:
a =
tan 9/2 (K-6)
so that the unknowns are now (xo, y , z , a, 3).
The quantitative measure of how well Eq. (K-5) approximates the measured
values of the axial coordinate, zm, is given by the coefficient of determination,
r2
N
N
Y(7 . _T \2 (K-7)
where zm is the average of the measured values of z. The numerator of the ratio
within the expression represents the variation in zm that is not accounted for
by Eq. (K-5) and the denominator represents the total variation in the data. The
difference between the denominator and numerator represents the variation in the
data that is accounted for by the working equation. (This difference is often
called the associated variation.) In other words:
r2 =
ASSOCIATED VARIATION
TOTAL VARIATION
Thus, the coefficient of determination is the ratio of the amount of variation in
the data accounted for by the working equation to the total variation in the data.
The objective of the analysis of the data was to maximize r2 which is equivalent
to minimizing the quantity:
184
-------
N
~j=l m' Cl (K-8)
Minimization of e requires differentiating Eq. (K-8) with respect to the unknowns
(x0, yQ, ZQ, a, 3), setting these five derivatives to zero, and solving the five
equations simultaneously. However, the system of equations is nonlinear (because
of the form of Eq. (K-7)). To avoid elaborate solution techniques, a parametric
approach was used in this study. The unknowns (xo, y0, g) were varied parametrically
and the unknowns (ZQ) a) were calculated from a set of linear equations. These
two equations are:
= 0 (K-9)
da
and
Using Eq. (K-7) in Eq. (K-8) and applying Eqs. (K-9) and (K-10) yields the two
equations:
N ,
(K-ll)
i = l
V / 1 (K-12)
-i = i ° ' 'J
where
has been introduced for convenience. These equations are linear in zo and a so
that:
N N N
a =
N N
i=i ' i=i
185
-------
and
(K-15)
o
For each set (x , y , z , a, 3), the coefficient of determination, r , was
calculated. The best fit of the data occurred for the highest value of r .
Of the five unknowns, four of these (XQ, y , ZQ, g) were used to transform
the raw data to the coordinate system in which the cone vertex lies at the origin
and the cone centerline lies on the z axis; data in the table in Appendix M con-
forms to this coordinate system. The remaining unknown, 9, is a property of the
injector spray. After determining the best set of (xo, y , z , 9, g), it was
necessary to estimate the precision with which the spray angle, 9, had been
determined.
It was assumed that (xo, y , z , 3) were fixed at the values giving the best
fit to the data. It was then assumed that a "spray" angle could be calculated
for each data point by applying Eq. (K-4) as follows:
N N
V V
2 -- i='Znni" ^ S'
-------
n-2
The results produced by Eq. (K-18) and Eq. (K-20) did not differ significantly.
Since each set of data consisted of a large number of measurements, the standard
error of estimate is an excellent approximation of the more familiar standard
deviation. Thus, approximately 95 percent of the data lies within two standard
deviations of the spray angle 6.
187
-------
APPENDIX L
GAS ANALYZER CALIBRATION PROCEDURES
The flame ionization detector output is related to the number of carbon atoms
released by the breaking of carbon-hydrogen bonds as the test gas passes through
the detector. By operating the analyzer at fixed temperature and pressure (400K,
1 atm) the output can be interpreted as being proportional to the mole fraction
of unburned hydrocarbon in the flame. For example, the mole fractions of fuel in
a stoichiometric mixture of propane and air is 0.04; because there are three car-
bon atoms/mole, the output of the analyzer in percent carbon (PC) for the mixture
would be 12. Likewise, for stoichiometric mixtures of high molecular weight fuels
of the form CnH2n would be 14.1
The gas analysis system was calibrated by passing gases of known composition
through the analyzer and comparing the percent carbon output to the values.
Various concentrations of methane, ethane, propane and butane were used in this
calibration procedure. The composition of the calibration gases, supplied by
Scott Research Laboratories, were guaranteed to within two percent of specifica-
tions. The calibration gases were injected into the previously evacuated gas
analysis stream upstream of the ten-port sampling valve as shown in Fig. L-l.
The calibration of the ionization gauge was carried out over five orders of
magnitude from 100 ppm methane to 100 percent butane. The resultant calibration
curve is shown in Fig. L-2. The abscissa is the percent carbon, PC, and the
ordinate is the integrated area (Ac) under the recorded concentration vs time
curves.
Expressions of the form
in PC--A + 'B in Ac (L_D
were fit to the data using a multiple linear regression analysis. The coefficients
were:
A = -3.12
B = .691
for .0001 < PC < .02
and
188
-------
HYDROCARBON ANALYZER CALIBRATION SYSTEM
EXHAUST
r
H2-N2
AIR
oo
eBRATION
GAS
TOTAL
VAPOR
IONIZATION
GAUGE
TEN PORT
SAMPLING
VALVE
-»- INPUT TO ELECTROMETER
ARGON
CARRIER
01
I
I
r
OVEN
P
I
-------
FIG. L-2
HYDROCARBON ANALYZER CALIBRATION CURVE
1II I I I I
S 0.09% METHANE , 99.91% N2
4.45% METHANE, 95.55% N2
0.0002
0.0001
0.01 0.025 0.05 0.1 0.2
0.5 1.0 2 5 10 20
PERCENT CARBON (PC)
50 100 200 500 1000
76-06-1 67-1
190
-------
A = -2.48
B = 1.165
for .0001 < C < 8
The standard deviation of the data about the curve is 0.116 in units of In PC.
191
-------
APPENDIX M
TABULATED HOLOGRAPHIC DATA
This appendix contains the following data for each of the holographic tests:
1) Spray surface coordinate data - see Fig. D-l for definition of coordinate
directions.
2) Spray thickness data - coordinates y^ and y~ give the locations of the
inner and outer surfaces of the hollow cone spray at the indicated values
of x and z.
3) Droplet diameter data - the number of droplets in a given size range
within a series of survey regions (see Fig. D-l) are tabulated. The
coordinates of the position of survey region are given in the table
of spray thickness data.
4) Reduced data - the Sauter mean diameter and spray thickness is reported
and the flow rate calculated from the droplet count, droplet size and
droplet velocity (from LV measurements) is compared with the metered fuel
flow rate.
192
-------
TABLE M-l
HOLOGRAPHIC RESULTS FOR RUN 129-7
Combusting Flow
Pressure (atm): 0.99
Swirl No.: 0.3
Equivalence Ratio: 0.66
Cone Angle: 69.4 + 2.0 deg
Fuel Type: Iso-octane
Inlet Temp. (K): 574
Fuel Flow (Kg/sec): 0.006
Airflow (Kg/sec): 0.137
o
Associated Data Variation, R :0.96
Point
x
Spray Surface Coordinates(cm)
z Point x
1
2
3
1
4
5
6
7
8
9
10
11
12
13
14
15
16
17
18
19
20
21
22
23
2k
25
26
27
28
29
30
31
.77
.77
.77
.77
77
.77
.77
.77
.77
.77
.77
.77
.77
.77
77
.77
.77
.77
77
.77
.77
.77
.77
.77
1.65
1.65
1.65
1.65
1.65
1.65
1.65
1.49
-1.14
1.22
-1.1*7
-2.06
-2.06
-1.97
-2.81
-2.81
-3.56
-3.67
-4.15
-4.15
-2.78
4.63
4.01
3.84
3.90
2.95
3.17
2.11
2.11
1.69
1.80
1.73
-1.1*0
-.90
1.1*5
1.45
-1.01
-2.25
2.06
1.97
2.05
1.96
3.14
3-14
3-15
4.21
4.21
5.29
5.28
6.36
6.36
6.4l
6.67
5.55
5.54
5-55
4.42
4.42
3.29
3.29
2.72
2.73
3.22
3.11
3.13
3.21
3.21
3.12
3.80
32
33
34
35
36
37
38
39
40
41
42
43
44 '
45
46
47
48
49
50
51
52
53
54
55
56
57
58
59
60
61
62
1.65
1.65
1.65
1.65
1.65
1.65
1.65
1.65
1.65
1.65
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
-.55
-.55
-.55
-.55
-.55
1.98
2.09
-2.19
2.14
2.88
2.88
-2.99
-2.99
-3.74
-3.79
1 -3.96
-4.01
-4.59
-4.92
-3.67
-3.6l
-2.92
-3.08
-2.39
-2.28
-1.82
-1.93
-1.52
-1.63
1.44 '
1.55
-4.14
-4.14
-3.72
-3.78
-2.86
3.94
3.95
3.80
3-95
4.91
4.91
4.70
4.70
5.72
5-72
5.71
5.71
6.51
6.50
5.28
5.28
4.21
4.21
3.13
3.14
2.60
2.60
1.90
1.90
2.00
1.90
6.25
6.25
5.28
5.28
4.21
193
-------
TABLE M-l. HOLOGRAPHIC RESULTS FOR RUN 129-7 (Cont'd)
v
Spray Surface Coordinates(cm)
Point x y z Point x y
63
6U
65
66
67
68
69
70
71
72
73
7U
75
76
77
78
79
80
81
82
83
8U
85
86
87
88
89
90
91
92
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
--55
-.55
-.55
-.55
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-2.81
-2.17
-2.11
-1.71
-1.82
-1.58
1.1+9
1.33
2.05
2.11
2.89
2.8U
3.62
. 3.73
U.Ul
U.U6
U.30
U.19
3.2U
3.57
2.U5
2.51
1.29
.37
.70
-.72
-.72
1.U5
-i.Uo
-i.Uo
U.21 93
3.1U 9^
3-1^ 95
2.61 96
2.60
2.06
2.17
2.16
3.29
3.29
U.Ul '
U.Ul
5.5^
5-5>+
6.66
6.66
6.U9
6.U9
5.52
5.5U
U.Uo
U.Uo
3.26
2.68
2.69
2.6U
2.6U
3.27
3.17
3.17
-1.65 -2.U2 U.23
-1.65 -2.U2 ^.23
-1.65 -3.3^ 5-29
-1.65 -3. to 5.29
194
-------
Pos.
1
2
3
H.
5
6
7
8
9
TABLE M-l. HOLOGRAPHIC RESULTS FOR ROT 129-7 (Cont'd)
Spray Thickness Data (cm)
1.6
1.3
.8
.U
.0
-.5
-1.1
-1.6
-2.2
2.0
2.1
2.2
2.3
2.5
2.1
2.1
i.U
.9
3.3 U.U
3.6 b.h
3-9 ^
3.5 ^
3.1 ^
3.3 ^
3.6 U.U
3.U U.^
2.3 U.U
Pos,
55
Spray Droplet Diameter Data (microns)
(Table entry is number of droplets observed)
83
111
139
167
195
222
250 278
1
2
3
1*
5
^
6
7
\
8
9
3
0
0
2
2
0
0
0
0
3
J+
3
8
k
7
5
5
5
2
2
0
0
0
1
2
1
5
0
1
1
0
0
1
0
2
0
0
0
1
0
0
0
0
1
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
Reduced Data
SMD
SMD
AVE. THICHUESS
AVE. Z
N-COUNTER
N-MEASURED
N-CALCULATED
AVE. THICMESS
STD. ERROR OF T
U.I MILS
105 MICRONS
1.2 CENTIMETERS
UA CENTIMETERS
58 6
15-7 x 10b
12.3 x 106
1.2 CENTIMETERS
.k CENTIMETERS
195
-------
TABLE M-2. HOLOGRAPHIC RESULTS FOR RUN 129-8
Non-combusting Flow Fuel Type: Iso-octane
Pressure (atm): 1.0 Inlet Temp. (K): 576
Swirl No.: 0.3 Fuel Flow (Kg/sec): 0.006
Equivalence Ratio: 0.66 Airflow (Kg/sec): 0.137
Q
Cone Angle: 81.4 + 2.6 Associated Data Variation, R^: 0.96
Spray Surface Coodinates(cm)
Point x y z Point x y z
1
2
3
1+
5
6
7
8
9
10
11
12
13
14
15
16
17
18
19
20
21
22
23
24
25
26
27
28
29
30
31
.11
.11
.11
.11
.11
.11
.11
.11
.11
.11
.11
.11
.11
.11
.11
.11
.11
.11
.11
.11
.11
.11
1.43
1.43
1.43
1.43
1.43
1.43
1.43
1.43
1.43
1.26
-1.48
1.26
-1.37
1.76
1.70
2.58
2.69
3.68
3.29
3.18
4.44
-4.55
-U.55
-3-95
-3.90
-2.96
-2.96
-1.76
-1.92
--93
--93
-1.10
-.88
-1.87
-2.14
-3.^0
-3.24
-4.50
3.29
3-18
l.4l
1.4l
l.4l
1.41
2.01
2.01
3.H
3.11
4.20
4.20
4.20
4.97
4.97
4.97
4.20
4.20
3.11
3.11
2.01
2.01
1.13
1.13
2.34
2.34
3.H
3.11
4.20
4.20
5.30
4.20
4.20
32
33
34
35
36
37
38
39
40
41
42
^3
44
1+5
46
47
48
U9
50
51
52
53
5^
55
56
57
58
1.43
1.43
1.43
1.43
1.43
1.43
1.15
-1.32
-1.32
-1.32
-1.32
-1.32
-1.32
-1.32
-1.32
-1.32
-1.32
-1.32
-1.32
-.88
-.88
-.88
-.88
-.88
-.88
-.88
-.88
4.17
4.17
2.14
2.14
-99
1.32
.22
2.30
2.25
2.91
2.63
3-73
3.73
^.39
-4.06
-2.85
-2.74
-1.97
-2.30
-1.76
1.8l
1.81
-1.21
-1.26
-1.26
-1.37
1.76
5-30
5.30
3.11
3.11
2.01
2.01
1.13
2.83
2.83
3.66
3.66
^.75
4.75
5.41
5.41
4.20
4.20
3.11
3.11
2.01
2.01
2.01
2.01
2.01
2.01
2.01
2.01
196
-------
Pos,
1
2
3
U
5
6
7
Pos.
TABLE M-2. HOLOGRAPHIC RESULTS FOR RIM 129-8 (Gont'd)
Spray Thickness Data (cm)
1.3
-9
A
.0
2.U
-1.0
-1.7
y
1.9
2.1
1.8
-.9
1.5
1.9
2.2
2.5
2.9
2.9
3.2
2.6
2.9
3-1
3.1
3.1
3.1
3.1
3.1
3-1
Spray Droplet Diameter Data (microns)
(Table entry is number of droplets observed)
55 83 111 139 167
195
222 250 278
1
2
3
k
5
6
7
0
0
0
0
0
0
0
0
0
0
6
6
0
5
0
1+
6
2
U
5
2
2
2
0
1
1
2
0
2
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
Reduced Data
SMD
SMD
AVE. THICKNESS
AVE. Z
N- COUNTED
N-MEASURED
N-CALCULATED
AVE. THICKNESS
STD. ERROR OF T
1+.7 MILS
118 MICRONS
,9 CENTIMETERS
3.1 CENTIMETERS
8.7
10
106
,9 CENTIMETERS
.3 CENTIMETERS
197
-------
TABLE M-3, HOLOGRAPHIC RESULTS FOR RUN 130-3
Non-combusting Flow
Pressure (atm): 1.0
Swirl No.: 0.3
Equivalence Ratio: 0.69
Cone Angle: 67,5 + 3.2
Fuel Type: Wo. 2 Fuel Oil
Inlet Temp. (K): 739
Fuel Flow (Kg/sec): 0.006
Airflow (Kg/sec): 0.137
2
Associated Data Variation, R : 0.94
Point
x
Spray Surface Coordinates(cm)
Point
x
y
1
2
3
4
5
6
7
8
9
10
11
12
13
Ik
15
16
17
18
19
20
21
22
23
2k
25
26
27
28
29
30
31
32
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
- = 55
-=55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
1.22
1-33
-.61
-.69
-.61
-1.29
-1.08
-1.08
1.62
1.51
2.02
2.18
2.18
2.85
2.69
2.80
-2.36
-2.22
-2.16
-1.45
-1.43
-2.81
-2.81
3.3k
3.56
3.3k
3.83
3.72
-2.75
-2.6k
-2.6k
-3.48
1.51
1.53
1.17
1.16
1.17
1.83
1.87
1.87
2.37
2.35
3.22
3-25
3.25
4.16
4.13
4.15
3.19
3.22
3.23
2.24
2.25
4.00
4.00
5.14
5.18
5.14
5.51
5.49
4.29
4.31
4.31
5.21
33
34
35
36
37
38
39
4o
41
42
43
44
45
46
47
48
49
50
51
52
53
54
55
56
57
58
59
60
61
62
63
64
-.55
-.55
-.55
-.55
.55
.55
= 55
.55
.55
.55
.55
55
.55
.55
= 55
.55
.55
.55
.55
.55
.55
.55
.55
.55
.55
.55
55
.55
.55
-1.54
-1.54
-1.54
-3-48
-3-48
-4.18
-4.18
-.09
-.20
-99
1.20
1.82
1.82
-1.15
-1.20
-1.20
-1.31
2.41
2.41
-2.12
-2.07
-2.23
-3.03
-2.97
-2.92
-3.17
-3-17
-3.13
3.97
2.75
2.75
1.73
-1.47
-1.20
-1.31
5.21
5.21
6.59
6.59
1.38
1.36
1.58
1.62
2.79
2.79
2.24
2.23
2.23
2.21
3.85
3.85
3.01
3.02
2.99
3.96
3.97
3.98
4.71
4.71
5-72
5.37
4.08
4.08
2.66
3.08
3.13
3.11
198
-------
TABLE M-3. HOLOGRAPHIC RESULTS FOR ROT 130-3 (Cont'd)
Spray Surface Coordinates(cm)
Point x y z Point x y
65
66
67
68
69
70
71
72
73
74
-1.54
-1.54
-1.54
-1.54
-1.54
-1.5^
-1.54
-1.54
-1.54
-1.54
1-77
1.88
2.69
2.80
-1.79
-1.84
-2.24
-2.24
3-53
3.69
3.68
3.70
4.74
4.76
3-91
3-90
5.16
5.16
6.23
6.26
199
-------
Pos.
TABLE M-3. HOLOGEAPHIG RESULTS FOR RIM 130-3 (Cont'd)
Spray Thickness Data (cm)
1 1.5 1.5 1-5 ^-2
2 l.l 1.3 1.3 ^.1
3 1.9 .2 .2 3-9
1* .2 2.2 2.2 I*. 3
5 -.2 2.1 2.1 l*.3
6 -.5 2.2 2.2 U.3
7 -.8 1.8 1.8 U.2
8 -1.1 1.5 1.5 ^.2
9 -l.U 1.8 1.8 U.2
Spray Droplet Diameter Data (microns)
(Table entry is number of droplets observed)
Pos. 55 83 111 139 167 195 222 250 278
1
2
3
h
5
6
7
8
9
0
0
0
0
0
0
0
0
0
3
0
0
0
0
0
0
0
0
2
1
h
0
3
3
5
3
2
1
1
3
5
1
1
0
0
0
0
2
0
0
1
1
0
1
1
0
0
0
1
1
0
0
0
0
0
0
0
0
0
1
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
Reduced Data
SMD 5.7 MILS
SMD Ikk MICRONS
AVE. THICKNESS .0 CENTIMETERS
AVE. Z k.2 CENTIMETERS
N-COUNTED ^0
N-MEASURED 3.8 x 106
N-CALCUIATED k.Q x 10^
AVE. THICKNESS .0 CENTIMETERS
STD. ERROR OF T .0 CENTIMETERS
200
-------
TABLE M-4. HOLOGRAPHIC RESULTS FOR RIM 130-1*
Combusting Flow-
Pressure (atm): 1.0
Swirl No.: 0.3
Equivalence Ratio: 0.66
Fuel Type: Ho. 2 Fuel Oil
Inlet Temp. (K): 758
Fuel Flow (Kg/sec): 0.006
Airflow (Kg/sec): 0.138
Cone Angle: 56.5 1 2.1 Associated Data Variation, R: 0.92
Spray Surface Coordinates(cm)
Point
y
Point
y
1
2
3
i
4
5
7
8
9
10
11
12
13
14
15
16
17
18
19
20
21
22
23
24
25
26
27
28
29
30
31
32
-.55
-.55
-.55
-.55
-^55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
-.55
--55
-.55
-.55
-.55
1.43
-1.34
1.35
-l.4o
-1.70
-1.65
1.67
1.67
-1.76
-2.18
-2.33
2.65
2.49
-2.45
-2.36
2.49
3.01
3.09
-2.63
-2.69
-2.79
-3.03
-3.19
-3.19
3.65
3.70
3-83
3.9^-
-3.57
-3.79
-3.84
-3.68
2.83
2.49
2.82
2.48
3.16
3.17
3.58
3.58
3.16
4.05
4.03
4.64
4.62
4.01
4.02
4.62
5.57
5.58
4.87
4.87
4.85
5.88
5.86
5.86
6.70
6.70
7.60
7.62
6.69
' 6.67
6.66
6.68
33
34
35
36
37
38
39
4o
4l
42
43
44
45
46
47
48
49
50
51
52
53
54
55
56
57
58
59
60
61
62
63
64
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
.77
.77
.77
.77
.77
.77
.77
.77
.77
.77
.77
77
.77
.77
.82
.58
.58
-.50
-.78
-.50
--99
-1.53
-1.64
1.25
1.35
1.19
1.85
1.85
1.74
-1.96
-1.75
-1.64
3.21
3.78
3.84
-3.26
-3.12
-2.78
-2.78
-2.83
-2.51
-2.05
-1.99
-1.99
-1.42
-1.37
3.48
3.^5
3.45
3.31
3.28
3.31
3.64
3.57
3.56
3.91
3.93
3-91
4.87
4.87
4.86
4.40
4.43
4.44
5.92
7.10
7.11
7.34
7.36
6.13
6.13
5.18
5.22
4.17
4.18
4.18
3.14
3.15
201
-------
TABLE E-k. HOLOGRAPHIC RESULTS FOR RUN 130-k (Cont'd)
Spray Surface Coordinates(cm)
Point x y z Point x y
65
66
67
68
69
70
71
72
73
7^
75
76
77
78
79
80
81
82
83
8U
85
86
87
88
89
90
91
.77
.77
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
-1.65
.77
.77
.77
.77
.77
77
.77
77
.77
77
.77
.77
-.67
-.72
-1.91
1.70
2.k6
2.52
-2.06
-2.11
-2.7k
-2.7k
3.1^
3-29
3.27
-3.15
-3.31
-1 .21
-1.21
1.97
1.89
2.17
2.17
2.17
2.7U
2.7k
3.32
3.32
-2.73
2. lt-1
2.1*0
k.kl
5.63
5-72
5-73
5.17
5.16
6.19
6.19
6.91
6.93
7.65
6.86
6.81*
2.62
2.62
3.01
3.00
k.03
k.03
k.03
k.Q2
k.82
5. 9k
5-9k
5.19
202
-------
TABLE M-4. HOLOGRAPHIC RESULTS FOR RUN 130-4 (Conf d)
Spray Thickness Data (cm)
x
2 T'l ° 2-9 4.7
3 i 2'6 2.0 47
0 I'l 3<1 "'8
5 ? 2'6 3.1 4.8
J --3 3.1
6 -.8 P.k
£i ^
9 -2.2 .5 -1.2
Spray Droplet Diameter Data (microns)
(Table entry is number of droplets observed)
3.0 4.8
.6 4.7
.2 4.7
Pos.
1
I,
<4
£
o
7
9
55
0
0
0
0
0
0
0
0
0
83
0
0
0
0
0
0
0
4
0
111
5
2
4
2
0
3
2
2
4
139
5
3
2
0
1
0
1
1
2
167
2
0
1
0
1
0
0
4
0
195
0
0
0
0
0
0
0
0
1
222
0
0
0
0
0
0
0
0
0
250
0
o
0
0
0
0
0
0
0
278
o
o
0
0
0
0
0
0
0
Reduced Data
SMD 5.4 MILS
SMD 137 MICRONS
AVE. THICKNESS .6 CENTIMETERS
AVE. Z 4.7 CENTIMETERS
H-COUNTED 46
N-MEASURED 3.2 x 1CT
N-CALCULATED 5.5 x 10°
AVE. THICKNESS .6 CENTIMETERS
STD. ERROR OF T .4 CENTIMETERS
203
-------
TABLE M-5. HOLOGRAPHIC RESULTS FOR RUN 130-7
Non-combusting Flow
Pressure (atm): 1.0
Swirl No.: 0.3
Equivalence Ratio: 0.67
Fuel Type: No. 2 Fuel Oil
Inlet Temp. (K): 523
Fuel Flow (Kg/sec): 0.0063
Airflow (Kg/sec): 0.137
Cone Angle: 75.1 + 3.2 Associated Data Variation, Rc: 0.95
Spray Surface Coordinates
Point
x
y
Point
x
y
1
2
3
1*
5
6
7
8
9
10
11
12
13
11*
15
16
17
18
19
20
21
22
23
21*
25
26
27
28
29
30
31
32
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.52
1.01
-.53
1.10
-.1*
1.21
1.1*3
-=99
1.57
-1.2U
1.1*0
-l.OU
-1.15
2.00
-1.65
2.30
-1.65
1.92
2.19
-2.19
2.36
-2.63
2.30
2.7l*
-3.13
2.71*
-2.7l*
-3.07
-3.02
3.2l*
2.7^
2.69
1.19
1.19
1.19
1.19
1.19
1.19
1.79
1.79
1.79
1.79
1.79
1.79
2.56
2.56
2.56
2.56
2.56
2.56
3.22
3.22
3-22
3.22
3.22
3.82
3.82
3.82
3.82
3.82
3.82
3-82
3.82
3.82
33
31+
35
36
37
38
39
1*0
1*1
1*2
U3
1*1*
U5
1*6
U7
1*8
U9
50
51
52
53
5^
55
56
57
58
59
60
61
62
63
61*
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-.63
-2.12
1.62
1-59
.66
.66
.66
.66
.66
.66
.66
.66
.66
.66
.66
.66
.66
.66
2.71*
3.68
3.59
3-62
3-57
-3.51
-3.1*0
-3.^9
-3.1*0
-U.06
-U.19
-U.lU
1*.17
U.17
1*.28
-1.51*
-1.5^
-1.5^
1.26
1.15
1.21
-1.5U
-1.87
-2.30
-2.il*
-2.19
1.59
1.70
1.65
2.52
2.36
2.52
3.82
^-75
U.75
l*-75
1*.75
U.75
l*-75
*4.75
U.75
5-^7
5.^7
5.^7
5.^7
5-^7
5.^7
3.00
3-00
3.00
1.79
1.79
1.79
1.79
1.79
2.50
2.50
2.50
2.50
2.50
2.50
3.1*1*
3.M*
3.M*
204
-------
TABLE M-5. HOLOGRAPHIC RESULTS FOR RUN 130-7 (Cont'd)
Spray Surface Coordinates
Point
Note:
130-7
x
65
66
67
68
69
70
71
72
73
74
75
76
77
78
79
80
81
82
83
84
85
86
87
88
89
90
.66
.66
.66
.66
.66
.66
.66
.66
.66
.66
.66
.66
.66
-1.54
-1.54
-1.54
-1.54
-1.54
-1.54
-1.54
-1.54
-1.54
-1.54
-1.54
-1.54
-1.54
2.36
-2.58
-2.50
-2.63
-2.63
2.63
2.4l
3.24
3.18
3.29
-3.40
-3.68
-3.51
.38
-1.21
-.66
-1.37
-1.32
1.26
1.21
1.43
-2.19
-2.19
-2.19
1.43
1.97
3.44
3.44
3.44
3.44
3.44
3.44
3.44
4.50
4.50
4.50
4.50
4.50
4.50
1.73
2.50
2.50
2.50
2.50
2.50
2,50
2.50
3.44
3.44
3.44
3.44
3.44
Point x y z
91 -1.54 2.03 3.44
92 -1.54 3.13 4.37
93 -1.54 3.18 4.37
94 -1.54 -2.96 4.37
95 -1.54 -3.07 4.37
Spray thickness and droplet size data were not available for Run
205
-------
TABLE M-6. HOLOGRAPHIC RESULTS FOR RUN 130-8
Combusting Flow
Pressure (atm): 1.0
Swirl No.: 0.3
Equivalence Ratio: 0.66
Cone Angle: 64.8 + 5.2
Fuel Type: Wo. 2 Fuel Oil
Inlet Temp. (K): 532
Fuel Flow (Kg/sec): 0.006
Airflow (Kg/sec): 0.137
2
Associated Data Variation, R : 0.93
Point
x
Spray Surface Coordinates
Point
x
y
1
2
3
4
5
6
7
8
9
10
11
12
13
14
15
16
17
18
19
20
21
22
23
2k
25
26
27
28
29
30
31
32
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
.00
,58
.44
.33
-.66
-.71
-.66
-1.32
-1.21
.66
.77
.66
.60
1.10
=93
93
-1.50
-1.5^
-1.76
-1.72
-1.72
1.68
1.54
1.97
2.08
-1.92
-1.97
-2.03
-2.63
-2.60
2.41
2.30
2.4l
.59
^59
.59
.59
59
.59
1.14
1.14
1.14
1.14
1.14
1.14
1.75
1.75
1-75
1.75
1.75
2.46
2.46
2.46
2.46
2.46
3.06
3.06
3.06
3.06
3.06
3.85
3.85
3.85
3.85
3.85
33
34
35
36
37
38
39
40
41
42
43
44
45
46
47
48
49
50
51
52
53
54
55
56
57
58
59
60
61
62
63
64
.00
.00
.00
.00
1.65
1.65
1.65
1.65
1.65
1.65
1.65
1.65
1.65
1.65
1.65
1.65
1.65
1.65
1.65
1.65
1.65
-1.32
-1.32
-1.32
-1.32
-1.32
-1.32
-1.32
-1.32
-1.32
-1.32
-1.32
-3.79
-3.62
3.29
3.29
3-29
3.0?
-2.30
-2.19
-1.43
-1.21
-1.32
-1.43
2.08
2.08
1.10
1.10
1.10
-.11
.99
-.33
-.33
.44
.44
-.44
-.44
1.26
1.43
1.18
-1.43
-1.32
1.8l
1.97
5.42
5.42
5.42
5.42
4-73
4.73
4.73
4.73
3.89
3.89
3.89
3.89
3.89
3.89
3.08
3.08
3.08
3.08
3.08
3.08
3.08
2.35
2.35
2.35
2~.35
3.10
3.10
3.10
3.10
3.10
3.89
3.89
206
-------
TABLE M-6. HOLOGRAPHIC RESULTS FOR RUN 130-8 (Cont'd)
Spray Surface Coordinates
Point x y z Point x y z
65 -1.32 1.87 3.89
66 -1.32 -1.97 3.89
67 -1.32 -1,97 3.89
68 -1.32 -2.U7 ^.98
69 -1.32 -2.U1 U. 98
70 -1.32 -2,58 U-98
71 -1.32 2.85 U.98
72 -1.32 3.18 IK 98
73 -1.32 2.85 l*.98
207
-------
Pos.
1
2
3
U
5
6
7
Pos.
TABLE M-6. HOLOGRAPHIC RESULTS FOR RUN 130-8 (Cont'd)
Spray Thickness Data (cm)
1.5
l.lf
.8
A
.0
-A
-1.3
yi
.2
.7
1.5
1.9
1.9
1.5
.5
.2
.7
1.5
1.9
1-9
1.5
.5
2.5
2.5
2.5
2.5
2.5
2 = 5
2.5
Spray Droplet Diameter Data (microns)
(Table entry is number of droplets observed)
83
111
139
167
195
222
250
278
1
2
3
k
5
6
7
0
0
0
0
0
0
0
0
0
1
0
0
0
0
3
6
6
6
U
k
0
u
1
1
8
1
U
U
1
1
1
1
0
2
0
0
0
1
0
0
0
0
0
0
0
0
0
0
1
0
0
0
0
0
0
0
0
0
0
0
0
0
0
Reduced Data
SMD
SMD
AVE. THICKNESS
AVE. Z
N-COUNTED
N-MEASURED
N-CALCULATED
AVE. THICKNESS
STD. ERROR OF T
5.5 MILS
139 MICRONS
.0 CENTIMETERS
2.5 CENTIMETERS
3.2 x 106
5.2 x 10°
.0 CENTIMETERS
.0 CENTIMETERS
208
-------
TABLE M-7. HOLOGRAPHIC RESULTS FOR RUN 132-4
Combusting Flow
Pressure (atm):
Swirl No.: 0.3
Equivalence Ratio: 0.64
Fuel Type: Iso-octane
Inlet Temp. (K): 528
Fuel Flow (Kg/sec): 0.006
Airflow (Kg/sec): 0.139
Cone Angle: 66.9 + 3-1 Associated Data Variation, R : 0-97
Spray Surface Coordinates
Point
x
Point
x
y
1
3
i
4
5
7
8
9
10
11
12
13
14
15
16
17
18
19
20
21
22
23
2k
25
26
27
28
29
30
31
32
-.11
-.11
-.11
-.11
-.11
-.11
-.11
-.11
-.11
-.11
-.11
-.11
-.11
-.11
-.11
-.11
-.11
-.11
-.11
-.11
-.11
-.11
-.11
-.11
-.11
1.26
1.26
1.26
1.26
1.26
1.26
1.26
1.65
1.70
2.20
3.12
3.18
4.05
3.88
4.45
-2.79
-2.84
-2.82
-2.49
-2.44
-2.05
-2.05
-1.56
-1.72
-1.28
-1.39
-.68
1.19
-.49
-.44
-93
93
.83
.88
1.71
1.71
1.88
2.80
2.85
2.51
2.52
3.51
4.71
4.71
5.91
5.89
6.83
4.25
4.25
4.08
3.56
3.57
3.06
3.06
2.56
2.54
2.04
2.03
1.55
1.63
.80
.80
^95
-95
2.42
2.43
3.45
3.45
3.47
4.67
4.68
33
34
35
36
37
38
39
40
4l
42
43
44
45
46
47
48
49
50
1.26
1.26
1.26
1.26
1.26
1.26
1.26
1.26
1.26
1.26
1.26
-1.76
-1.76
-1.76
-1.76
-1.76
-1.76
-1.76
3.56
3.28
4.37
-2.75
-2.75
-3-59
-3.48
-3.48
-1.66
-1.66
-l.ll
-.46
-l.4o
-2.34
-2.91
-3.36
3.72
3.45
5.86
5.83
s v ^ ,_/
7.04
4.47
i |
4.47
5.10
5.12
5.12
3.54
3.54
2.50
2.62
3-13
4.24
5.4o
6.07
6.87
5.84
209
-------
Pos,
1
2
3
4
5
6
7
8
9
Pos.
TABLE M-7. HOLOGRAPHIC RESULTS FOE RUN 132-4 (Cont'd)
Spray Thickness Data (cm)
55
1.7
1.1*
1.2
.5
-.2
-.7
-1.3
-2.0
-2.6
2.0
1.9
2.1
2.1
2.6
2.3
2.3
1.4
1.0
3.0
3.2
3.2
3.7
3.5
3.9
3.8
3.1
2.4
4.6
4.6
4.7
4.7
4.7
4.7
4.7
4.6
4.6
Spray Droplet Diameter Data (microns)
(Table entry is number of droplets observed)
83
ill 139
167
195
Reduced Data
SMD
SMD
AVE. THICKNESS
AVE. Z
N-COUNTED
N-MEASURED
N-CALCULATED
AVE. THICKNESS
STD. ERROR OF T
4.1 MILS
104 MICRONS
1.2 CENTIMETERS
4-7 CENTIMETERS
57 (,
5.1 x 10°
12.5 x 106
1.2 CENTIMETERS
.3 CENTIMETERS
250
278
1
2
3
4
5
6
7
8
9
6
0
0
0
0
4
0
7
7
1
5
0
3
7
3
6
3
2
0
3
3
4
l
0
6
0
i
i
0
0
0
1
0
0
0
0
0
1
1
1
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
210
-------
TABLE M-8. HOLOGMPHIC RESULTS FOR RUN 133-1
Non-combusting Flow
Pressure (atm): 1.0
Swirl No.: 0.6
Equivalence Ratio: 0.67
Fuel Type: Iso-Octane
Inlet Temp. (K): 54l
Fuel Flow (Kg/sec); 0.006
Airflow (Kg/sec): 0.137
Cone Angle: 70.5 + 2.4 Associated Data Variation, R2: 0.97
Spray Surface Coordinates
Point
x
y
Point
x
1
2
3
4
5
6
7
8
9
10
11
12
13
14
15
16
17
18
19
20
21
22
23
2k
25
26
27
28
29
30
31
J-*-
32
-.22
-.22
-.22
-.22
-.22
-.22
-.22
-.22
-.22
-.22
-.22
-.22
-.22
-.22
-.22
-.22
-.22
-.22
-.22
-.22
-.22
-.22
-.22
-.22
-.22
1.15
1.15
1.15
1.15
1.15
1.15
1.15
1.07
-.84
.86
-.89
.68
-.47
-1.40
-1.29
-2.13
-2.13
-2.59
-3.05
-3.68
-3.81+
-4.88
4.01
4.50
3,52
3.04
2.12
2.01
1.80
1.86
i.4o
1.59
-.87
-1.18
-1.48
-2.32
-3-11
-3.46
-3^90
i;4i
1.24
1.39
1.24
.88
.78
2.02
2.03
2.84
2.84
3.68
4.46
5.^0
5.38
6.51
5.63
6.28
M3
U.17
3.26
3.25
2.52
2,52
1.82
1.84
1.62
2.0U
2.89
3-70
4.51
5.36
6.04
33
34
35
36
37
38
39
40
41
42
43
44
45
46
47
48
49
50
51
52
53
1.15
1.15
1.15
1.15
1.15
1.15
1.15
1.15
1.15
-1.59
-1.59
-1.59
-1.59
-1.59
-1.59
-1.59
-1.59
-1.59
-1.59
-1.59
-1.59
-4.34
3.69
3.63
3.83
2.90
2.24
2.08
l.6l
1.14
-1.30
.09
-.13
-1.88
-2.94
-3.68'
-4.08
4.18
3.54
2.31
1.29
.12
6.66
5-55
5.54
6.39
4.49
3.77
3.75
2.89
1.91
3.35
2.59
2.57
3.74
4.53
5.34
6.19
6.25
4.65
3.66
2.75
2.26
211
-------
Pos,
1
2
3
4
5
6
7
Pos,
TABLE M-8. HOLOGRAPHIC RESULTS FOR RUN 133-1 (Cont'd)
Spray Thickness Data (cm)
x
1.3
.8
.3
-.1
-.5
-i.o
-1.5
1.1
1.4
1.5
1.4
1.3
.7
-.6
2.4
2.2
2.7
2.6
2.6
2.6
1.8
3-2
3.2
3.3
3.2
3.2
3-2
3.1
Spray Droplet Diameter Data (microns)
(Table entry is number of droplets observed)
55
83
111
139
167
195
Reduced Data
SMD
SMD
AVE. THICKNESS
AVE. Z
N-COUNTED
N-MEASURED
N-CALCULATED
AVE. THICKNESS
SID. ERROR OF T
5.0 MILS
128 MICRONS
1.1 CENTIMETERS
3.2 CENTIMETERS
46
4.4
6.8
x io
x io
1.1 CENTIMETERS
.3 CENTIMETERS
250
278
1
2
3
4
5
6
7
0
0
0
0
0
0
0
0
0
3
0
1
3
0
5
2
1
3
3
9
5
2
2
0
2
1
2
3
2
2
0
0
0
1
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
212
-------
TABLE M-9. HOLOGRAPHIC RESULTS FOR ROT 133-3
Combusting Flow
Pressure (atm): 1.0
Swirl Wo.: 0.6
Equivalence Ratio: 0.66
Fuel Type: Iso-Octane
Inlet Temp. (K): 537
Fuel Flow (Kg/sec): 0.006
Airflow (Kg/sec): 0.136
2
Cone Angle: 63.8 + 3.0 Associated Data Variation, R : 0.94
Spray Surface Coordinates
Point
y
Point
1
2
3
i.
M-
5
f.
D
9
10
11
12
13
_ i
14
15
16
17
18
19
20
21
22
23
24
25
26
27
28
29
30
31.
32
.16
.16
.16
_ /
.16
.16
_ /
.16
.16
.16
.16
.16
.16
.16
.16
.16
.16
.16
.16
.16
.16
.16
1.54
1.54
1.54
1.54
1.54
1.54
1.54
1.54
1.54
1.54
1.5^
1.&
-1.65
-99
-99
= 55
1.^3
1.92
2.52
3.13
3.62
3.8U
-3 ..51
-k.U.
-^.39
-3.2J+
-3.07
-2.52
-2.08
-1.87
-1.21
-.77
-.22
.88
2.08
2.52
2.91
3.07
2.91
3.62
-4.06
-3.13
-2. hi
-1.87
1.95
1.95
1.95
1.02
2.33
3.21
3.87
4.86
5.62
5.90
5.90
6.72
7.27
5.62
4.97
k.lk
3-27
2.33
1.89
1.02
2.22
2.94
3.43
4.31
5.19
6.06
6.06
6-94
6.94
5-95
5.24
4.53
33
34
35
36
37
38
39
40
4l
42
43
44
45
46
47
48
49
50
1.54
-1.21
-1.21
-1.21
-1.21
-1.21
-1.21
-1.21
-1.21
-1.21
-1.21
-1.21
-1.21
-1.21
-1.21
-1.21
-1.21
-1.21
-1.04
.33
.88
1.43
*j
2.03
2.30
3.07
2.80
4.17
-4.33
-4.66
-3.51
-2.85
-2.52
-1.97
-1.48
-1.15
.49
3.70
1.89
2.44
3.21
j # t .1
3.87
4.53
5.08
6.72
6.72
6.72
7-49
6.01
5.35
4.53
3.76
2.99
2.50
1.95
213
-------
TABLE M-9- HOLOGRAPHIC RESULTS FOR RIM 133-3 (Cont'd)
Spray Thickness Data (cm)
Pos,
1
2
3
U
X
-.1
-.5
-1.2
-1.8
.9
-9
2.0
2.1
1.3
2.0
3.7
3.7
3.7
3.7
Pos.
1
2
3
55
0
0
Spray Droplet Diameter Data (microns)
(Table entry is number of droplets observed)
83
2
2
3
o
ill 139 167
2
k
1
7
2
2
1
2
0
0
0
1
195
0
0
0
0
222
0
0
0
0
250 278
0
0
0
0
0
0
0
0
Reduced Data
SMD
SMD
AVE. THICKNESS
AVE. Z
N-COUNTED
N-MEASURED
W-CALCULATED
AVE. THICKNESS
STD. ERROR OF T
U.7 MILS
119 MICRONS
-9 CENTIMETERS
3.7 CENTIMETERS
27 *
k.kx 106
8.3 x 106
.9 CENTIMETERS
.3 CENTIMETERS
214
-------
REFERENCES
1. Norster, E. R. and A. H. Lefebvre: Effects of Fuel Injection Method on Gas
Turbine Combustor Emissions. Emissions from Continuous Combustion Systems,
Cornelius, W. and W. G. Agnew (eds.), New York, Plenum Press, pp. 255-278,'
j.y / <
2. Grobman, J. S.: Effect of Operating Variables of Pollutant Emissions from
Aircraft Turbine Engine Combustors. Emissions from Continuous Combustion
Systems, Cornelius, W. and W. G. Agnew (eds.), New York, Plenum Press, pp. 279-
303, 1972.
3. Tuttle, J. H., R. A. Altenkirch and A. M. Mellor: Emissions ^rom and Within
an Allison J-33 Combustor. II. The Effect of Inlet Air Temperature. Comb.
Sci. Technol. ]_: 125-134, 1973.
4. Pompei, F. and J. B. Heywood: The Role of Mixing in Burner-Generated Carbon
Monoxide and Nitric Oxide. Comb. Flame 19; 407-418,. 1972.
5. Mellor, A. M.: Simplified Physical Model of Spray Combustion in a Gas Turbine
Engine. Comb. Sci. Technol. 8_: 101-109, 1973.
6. Bowman, C. T. and L. S. Cohen: Influence of Aerodynamic Phenomena on Pollutant
Formation in Combustion. Environmental Protection Agency, Research Triangle
Park, N. C., Publication Number 650/2-75=061a, p. 159, July 1975.
7. Tuttle, J. H. , M. B. Colket, R. W. Bilger and A. M. Mellor: Characteristic
Times for Combustion and Pollutant Formation in Spray Combustion. Paper
presented at the 16th Symposium (International) on Combustion. Cambridge,
Mass., August 1976.
8. Appleton, J. P. and J. B. Heywood: The Effects of Imperfect Fuel-Air Mixing
in a Burner on NO Formation from Nitrogen in the Air and the Fuel. Fourteenth
Symposium (International) on Combustion. Pittsburgh, PA. The Combustion
Institute, pp. 77-786, 1973.
9. Spadaccini, L. J., F. K. Owen and C. T. Bowman: Influence of Aerodynamic
Phenomena on Pollutant Formation in Combustion of Gaseous Fuels. Environ-
mental Protection Agency, Research Triangle Park, NC, Publication Number
600/2-76-247a, September 1976.
10. Kerr, N. M. and D. Fraser: Swirl. Part I, Effect on Axisymmetrical
Turbulent Jets. J. Inst. Fuel 38; 519:538, 1965.
215
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11. Bowman, C. T. and L. S. Cohen: Influence of Aerodynamic Phenomena on
Pollutant Formation in Combustion. Environmental Protection Agency,
Research Triangle Park, NC, Publication Number EPA 650/2-75-061a, July
1975.
12. Tuttle, J. H., R. A. Altenkirch and A. M. Mellor: Emissions From an Within
an Allison J-33 Combustor II. The Effect of Inlet Air Temperature. Comb.
Sci. Technol. ]_: 125-134, 1973.
13. Beer, T. M. and N. A. Chigier: Combustion Aerodynamics, J. Wiley & Sons,
p. 142 (1972).
14. Tuttle, J. H., R. A. Shisler and A. M. Mellor: Nitrogen Dioxide Formation
in Gas Turbine Engines. Measurements and Measurement Methods. Comb. Sci.
Technol. 9: 261-271, 1975.
15. Wadleigh, R. R. and R. A. Oman: Instrumentation to Measure Composition and
Temperature of High-Velocity, Two-Phase, Two-Component Flows. MIT Dept.
of Mech. Engrg. Gas Turbine Laboratory, July 1956.
16. Kennedy, J. B. and J. B. McVey: An Experimental Study of Fuel Spray
Vaporization Rates. United Technologies Research Center Report UAR-N139,
September 27, 1974 and Eleventh JANNAF Combustion Meeting, Vol. II, CPA,
Publication Number 261, pp. 387-406, December 1974.
17. Owen, F. K.: Laser Velocimeter Measurements of a Confined Turbulent Diffusion
Flame Burner. United Technologies Research Center, Fourteenth AIAA Aerospace
Sciences Meeting, Washington, D.C., p.10, 1976.
18. Lindgren, B. W. and G. W. McAlrath: Introduction to Probability and
Statistics, New York, Macmillan, p. 165, 1959.
19- Becker, H. A., H. C. Hottel and G- C. Williams: On the Light-Scatter Technique
for the Study of Turbulence and Mixing. J. Fluid Mech. 30: 259-284, 1967.
216
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NOMENCLATURE
2
A = Area, cm
d = Outer diameter or air annulus, cm, or
Peak to peak fringe spacing
d^ = Inner diameter of air annulus, cm
d^ = Measured droplet diameter, microns
dm = Calculated mean droplet diameter, microns
D = Combustor diameter, m
D = Particle diameter
D = Spray diameter, cm
f = Frequency, Hz, or
Lens focal length
f-r. = Doppler frequency, Hz
fQ = Offset frequency, Hz
K = Cunningham constant 1.8
K = Defined by Appendix A
£ = Mean free path, cm
L = Height of air annulus, cm
M = Mass flow rate, kg/sec
n. = Number of droplets of diameter, d^
N = Total number of samples
P = Error as defined by Eqs. (14) and (15)
r2 = Coefficient of determination as defined by Eq. (32)
R = Radius, m
217
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NOMENCLATURE (CONT'D)
R = Combustor radius, m
S = Swirl number as defined by Eq. (1), or
Displacement
SMD = Sauter Mean Diameter
S = Calculated variance in the variable x as defined in Eq. (17)
X
t = Spray thickness, cm
T = Temperature, °K
u = Rms particle velocity, m/sec
U = Convective velocity, m/sec
U. = Instantaneous axial velocity, m/sec
u' = Axial velocity fluctuation, m/sec
V = Droplet velocity, m/sec
V f = Combustor reference velocity, m/sec
W = Mean tangential gas velocity, m/sec
w' = Tangential velocity fluctuation, m/sec
Wf = Fuel flow rate, kg/sec
x = Random variable
X = Axial distance, m
Z = Swirler hub-to-tip ratio
Z = Axial coordinate
m
3 = True mean of the variable x, or
Angle of rotation
Y = Directional intermittency
218
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NOMENCLATURE (CONT'D)
Y = Error in the mean as defined by Eq. (14)
YT = Error in the variance as defined by Eq. (15)
TI = Swirl vane angle, deg
6 = Angle, deg
X = Wavelength, m
y = Viscosity, gm/cm-sec
p = Density, gm/cc
p = Particle density, gm/cc
cr = Rms velocity, m/sec
a. = Error of estimate as defined by Eq. (20)
9
a~ = Effor of estimate as defined by Eq. (18)
6
$ = Overall fuel-air equivalence ratio = (mfue]/mair)/(mfuei/mair)stoich
219
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TECHNICAL REPORT DATA
(Please read Instructions on the reverse before completing)
1. REPORT NO.
EPA-600/7-79-003
2.
3. RECIPIENT'S ACCESSION-NO.
4. TITLE AND SUBTITLE
Influence of Aerodynamic Phenomena on Pollutant
Formation in Combustion (Phase n. Liquid Fuels)
5. REPORT DATE
January 1979
6. PERFORMING ORGANIZATION CODE
7. AUTHOR(S)
L.J.Spadac cirri, John McVey, Jan Kennedy,
A.S.Kesten, F.K.Owen, and C.T. Bowman
8. PERFORMING ORGANIZATION REPORT NO
9. PERFORMING ORGANIZATION NAME AND ADDRESS
United Technologies Research Center
300 Main Street
East Hartford, Connecticut 06108
10. PROGRAM ELEMENT NO.
EHE624A
11. CONTRACT/GRANT NO.
68-02-1873
12. SPONSORING AGENCY NAME AND ADDRESS
EPA, Office of Research and Development
Industrial Environmental Research Laboratory
Research Triangle Park, NC 27711
13. TYPE OF REPORT AND PERIOD COVERED
Final: 8/76 - 10/77
14. SPONSORING AGENCY CODE
EPA/600/13
is. SUPPLEMENTARY NOTES
project officer is W. Steven Lanier, MD-65, 919/541-
2432. E PA- 600/2 -76-247a was the Phase I report.
16. ABSTRACT The reporj. gjves results of an experimental investigation of the effects of
the interaction between physical and chemical processes on pollutant formation and
destruction in a liquid-fuel, turbulent-diffusion flame burner. The effects of fuel
type, inlet air swirl, inlet air temperature, and combustor pressure on the spray
characteristics and the time-mean and fluctuating flow field structure were deter-
mined, using probing and optical techniques. Changes in the spray and flow field
structure were correlated with changes in pollutant emissions from the burner. The
investigation showed that varying these operating parameters produces major chan-
ges in spray dynamics, vaporization rates, and time-averaged fuel/air distribution
within the burner which significantly influence energy release rates and pollutant
formation and destruction. Significant differences were found between the mean
velocities of the gas and fuel droplets which likely influence droplet vaporization
rates and mixing of the vaporized fuel and air.
17.
KEY WORDS AND DOCUMENT ANALYSIS
DESCRIPTORS
b.lDENTIFIERS/OPEN ENDED TERMS
c. COSATI Field/Group
Air Pollution
Nitrogen Oxides
Carbon Monoxide
Oxygen
Propane
Fuel Oil
Speed Indicators
Flames
Kinetics
Aerodynamics
Spraying
Lasers
Holography
Air Pollution Control
Stationary Sources
Liquid Fuels
No. 2 Fuel Oil
Iso Octane
13 B
07B
07C
21D
14 B
2 IB
2 OK
20D
07A
20E
8. DISTRIBUTION STATEMENT
Unlimited
19. SECURITY CLASS (This Report)
Unclassified
21. NO. OF PAGES
20. SECURITY CLASS (Thispage)
Unclassified
22. PRICE
EPA Form 2220-1 (9-73)
220
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