RESEARCH REPORT
A STUDY OF THE INFLUENCE OF FUEL ATOMIZATION
VAPORIZATION, AND MIXING PROCESSES ON POLLUTANT
EMISSIONS FROM MOTOR-VEHICLE POWERPLANTS
to
ENVIRONMENTAL PROTECTION AGENCY
OFFICE OF AIR PROGRAMS
Contract No. CPA 70-20
January 31, 1972
€*Batteiie
Columbus Laboratories
-------
~-----------~
,
~ --~--
----,-- ----- -- - --~._--_.__.-
BATTELLE'S COLUMBUS LABORATORIES comprises the origi-
nal research center of an international organization devoted to research
and development.
Battelle is frequently described as a "bridge" between science and
industry - a role it has performed in more than 90 countries. It
conducts research encompassing virtually all facets of science and its
application. It also undertakes programs in fundamental research and
education.
Battelle-Columbus - with its staff of 2500 - serves industry and
government through contract research. It pursues:
. research embracing the physical and I ife sciences, engi-
neering, and selected social sciences
. design and development of materials, products, processes,
and systems
. information analysis, socioeconomic and technical eco-
nomic studies, and management planning research.
505 KING AVENUE. COLUMBUS, OHIO 43201
-------
PHASE II REPORT
on
A STUDY OF THE INFLUENCE OF FUEL ATOMIZATION
VAPORIZATION, AND MIXING PROCESSES ON POLLUTANT
8~ISSIONS FROM MOTOR-VEHICLE POWERPLANTS
to
ENVIRONMENTAL PROTECTION AGENCY
OFFICE OF AIR PROGRAMS
Contract No. CPA 70-20
January 31, .1972
by
D. A. Trayser, J. A. Gieseke, R. D. Fischer, and
F. A. Creswick
BATTELLE
Columbus Laboratories
505 King Avenue
Columbus, Ohio 43201
Franklin County
-------
ABSTRACT
This report summarizes the Phase II study of automotive-engine
induction systems conducted by Battelle-Columbus for the Environmental Pro-
tection Agency ~nder Contract Number CPA-lO-20. Results of the Phase I study
(Under Contract Number CPA 22-69-9) are covered in a report dated April 30,
1969.
The objectives of this experimental Phase II program were to obtain
experimental data on droplet impaction characteristics, fuel-film fluw vd
manifold walls, and fuel vaporization for a better understanding of induction-
system phenomena; and to demonstrate the potential of improved fuel atomization,
fuel vaporization, and intake manifold design for improving air-fuel mixing and
distribution.
Results of this study showed that ultrafine atomization, minimum
manifold-passage turning angle, long passage bend radii, and low air velocity
can reduce droplet impaction in an induction system.
However, appreciable
droplet impaction still occurs, even with droplet sizes as low as 14 microns,
because of deposition by flow-induced air turbulence.
Consequently, fuel
vaporization by intake air preheating and by manifold surface heating is recom-
mended in addition to improved atomization to reduce further the fuel film on
the wall.
An improved induction-system concept recommended by the project staff
includes: an air-atomizing fuel nozzle operating on the ultrasonic Hartmann-
whistle principle, a mixing section similar in construction to a can-type com-
bustor, an unconventional throttle valve such as a plug valve between the
mixture
mixture
generator and the intake manifold or a butterfly valve upstream of the
generator, and a single-plane manifold with a separate branch for each
having equally spaced, equal-area, radial entrances.
Recommendations for additional research include: experimental studies
cylinder
of fuel-droplet transport
design and development of
from this Phase II study.
and vaporization, with emphasis on vaporization, and
the advanced induction-system concepts resulting
-------
TABLE OF CONTENTS
INTRODUCTION.
. . . . .
. . . . .
. . . . .
. . . .
. . . .
OBJECTIVES. .
. . . .
. . . . .
. . . . . .
. . . .
. . . .
SUMMARY OF STUDY.
. . . . . . . . . . . .
. . . . . . .
. . . . .
CONCLUSIONS
. . . . . . .
. . . . . . . . .
. . . . . .
. . . . .
RECOMMENDATIONS
. . . . . .
. . . . . . . . . .
. . . . . .
EXPERIMENTAL WORK
. . . . . . . . .
. . . .
. . . . . .
. . . . .
The Laboratory Flow System
Manifold Test Section Geometries
Discussion of Fluids
Mixture Quality Measurement. . . . . .
Experimental Program. . . . . . . . .
Experimental Results . . . . . .
. . . . .
. . . . . .
. . . .
. . . .
. . . . . . . .
. . . . .
. . . . . .
. . . .
. . . .
. . . . . . . .
DESIGN CONCEPTS
. . . . .
. . . .
. . . .
Induction System Concept. . . . . . . . . . . .
Fuel Vaporization. . . . . . . . . . . . . . . . . . .
ACKNOWLEDGEMENTS.
........
. . . . . .
. . . . . . .
LIST OF REFERENCES. . .
. . . . . . . . . .
. . . . . .
. . . . .
APPENDIX A, IMPACTOR STAGE DESIGN FOR VAPOR DROPLET SAMPLERS. . .
APPENDIX B, CALIBRATION CURVES. . .
. . . . . . . . . . . .
LIST OF FIGURES
FIGURE 1.
Schematic of Laboratory Flow System for Induction
System Fundamental Studies. . . . . . . . . . .
FIGURE 2.
Oscilloscope Traces of Pulsating Flow. . . .
. . . . .
FIGURE 3.
Time Variation of Pulsating Flow
. . . . .
FIGURE 4.
Air Flow Capacity of Laboratory Flow System.
. . . . .
FIGURE 5.
Layout Drawing of Mixture Generator. . . . .
. . . . .
FIGURE 6.
Detail of Mixture Generator Fuel Skimmer System.
Page
1
2
2
4
6
7
7
22
30
35
48
56
83
83
98
104
105
A-I
B-1
9
11
12
14
16
18
-------
LIST OF FIGURES (Continued
Page
FIGURE 7.
Modification to Mixture Generator to Reduce Impaction. .
21
FIGURE 8.
Detail of Discharge Burner
. . . .
. . . .
. . . . . .
23
FIGURE 9A. 30-Degree-Bend Manifold Test Section . . . . . . . . . . 26
FIGURE 9B. 90-Degree, O-Inch-Radius-Bend Manifold Test Section. . 27
FIGURE 9C. 90-Degree, 4-Inch-Radius-Bend Manifold Test Section. . . 28
FIGURE 10. Sketch of Aluminum Instrumented Test Section Showing
Thermocouple Locations. . . . . ... ... . . . . .
29
FIGURE 11. Comparison of Liquid Viscosities
. . . . .
. . . . . . .
32
FIGURE 12. Standard Distillation Curves for a Representative
Composite Gasoline and for stoddard Solvent. . . . . .
34
FIGURE 13. Vapor/Droplet Sampling Probe -- Design 1
. . . . . . . .
37
FIGURE 14. Vapor/Droplet Sampling Probe -- Design 2
. . . .
. . . .
38
FIGURE 15. Vapor/Droplet Sampling Probe -- Design 3 . . . . . 38
FIGURE 16. Simple Tube Sampling Probe . . . . . . . . 40
FIGURE 17. Fuel-Film Wall Skimmer Design . . . . . . . . . . 41
FIGURE 18. Equilibrium Vapor Content in Air Over Stoddard
Solvent at Atmospheric Pressure. . . . . . . . . . . .
44
FIGURE 19. Flow Diagram of Sampling System
. . . . . . .
. . . . .
46
FIGURE 20. Cross-Section of Manifold Test Section Showing
Probe Positions as Viewed from Outlet. . . . . . . . .
50
FIGURE 21. Collection Apparatus for Drop Sizing
. . . . .
. . . .
53
FIGURE 22. Size Distributions for Drops Entering Test Sections
55
FIGURE 23. Wall Fuel-Film Flow Rate as a Function of Air Flow
Rate and Test-Section Geometry . . . . . . . . .
69
FIGURE 24. Fraction of Liquid Fuel Impacted as a Function of
Air Flow Rate and Test-Section Geometry . . . . .
70
FIGURE 25. Entrained-Fue1-Drop1et Stratification at Straight
Test-Section Exit. . . . . . . . . . . . . . . .
72
-------
FIGURE 26.
FIGURE 27.
FIGURE 28.
FIGURE 29.
FIGURE 30.
FIGURE 31.
FIGURE 32.
FIGURE 33.
FIGURE 34.
FIGURE 35.
FIGURE 36.
FIGURE 37.
FIGURE 38.
FIGURE 39.
FIGURE 40.
FIGURE 41.
FIGURE 42.
LIST OF FIGURES (Continued)
Entrained-Fuel-Droplet Stratification at
30-Degree-Bend Test-Section Exit. . . .
. . . . . . . .
Entrained-Fuel-Droplet Stratification at 90-Degree,
4-Inch-Radius-Bend Test-Section Exit. . . . . . . . . .
Entrained-Fuel-Droplet Stratification at 90-Degree,
O-Inch-Radius-Bend Test-Section Exit. . . . . . .
Wall Fuel-Film Flow Patterns in 30-Degree Test Section.
Wall Fuel-Film Flow Patterns in .90-Degree,
4-Inch Radius Test Section. . . . . . . .
. . . . . . .
Wall Fuel-Film Flow Patterns in 90-Degree,
O-Inch Radius Test Section. . . . . . . .
. . . . . . .
Sketch Layout of Prototype Improved Induction System--
Section View. . . . . . . . . . . . . . . . . . . . . .
Sketch Layout of Prototype Improved Indu~tion System--
Plan View . . . . . . . . . . . . . . . . .
Sketch Layout of Intake-Passage Configuration for
Paired-Port Engine -- Section View. . . . . . . .
Sketch Layout of Intake-Passage Configuration for
Paired-Port Engine -- Plan View. . . . . . . . . .
Sketch of Proposed Air-Atomizing Hartmann-Whistle-
Type Fuel Nozzle. . . . . . . . . . . . . . . . .
Sketch of Proposed Induction System With Plug Throttle.
Sketch Layout of Proposed Induction System with
Alternative Inlet-Air Throttle -- Section View.
. . . .
Sketch Layout of Proposed Induction System with
Alternative Inlet-Air Throttle -- Plan View. . . . . .
Fin and Tube Geometry for Fuel Vaporization Chamber
Sketch Layout of Vaporization Chamber
. . . . .
Fuel Vaporization Apparatus
. . . . . .
. . ., .
Page
72
73
73
76
77
78
84
85
86
87
90
93
95
96
. . 101
. 103
. 104
-------
FIGURE A-!.
FIGURE A-2.
FIGURE A-3.
FIGURE B-1.
FIGURE B-2.
FIGURE B-3.
FIGURE B-4.
FIGURE B-5.
FIGURE B-6.
FIGURE B-7.
FIGURE B-8.
FIGURE B-9.
LIST OF FIGURES (Continued)
Droplet Size Distribution Used for Calibration
of the Vapor/Droplet Sampler. . . . . . . . . .
Vapor/Droplet Sampler Calibration Arrangement
. . . . .
Collection Efficiency of the Vapor/Droplet
Sampler as a Function of Droplet Size. . .
. . . . . .
Computed Flow vs Pressure Differential Data for
Square-Edged Orifice Flowmeter with 1 D and 1/2 D Taps,
Calibration Curve for Fuel System Flow Meter
. . . . .
Drop Size Distribution at 28,000 RPM and 1 cm3/sec
Drop Size Distribution at 40,000 RPM and 1 cm3/sec
Drop Size Distribution at 46,000 RPM and 1 cm3/sec
Drop Size Distribution at 53,000 RPM and 1 cm3/sec
Drop Size Distribution at 59,000 RPM and 1 cm3/sec
3
Drop Size Distribution at 64,000 RPM and 1 cm /sec
3
Drop Size Distribution at 32,000 RPM and 0.5 cm /sec
3
FIGURE B-10. Drop Size Distribution at 43,000 RPM and 0.5 cm /sec
50,000 RPM and 3
FIGURE B-1!. Drop Size Distribution at 0.5 cm /sec
B-12. RPM and 3
FIGURE Drop Size Distribution at 56,000 0.5 cm /sec
3
FIGURE B-13. Drop Size Distribution at 61,000 RPM and 0.5 cm /sec
FIGURE B-14. 3
Drop Size Distribution at 65,000 RPM and 0.5 cm /sec
FIGURE B-15. Standard Calibration of Primary Sampling Rotameters . .
FIGURE B-16. Temperature and Pressure Corrections for
Sampling Rotameters . . . . . . . . . . . . . . .
FIGURE B-17. Hydrocarbon Analyzer Calibration
. . . . . . . .
Page
A-3
A-4
A-6
B-2
B-3
B-4
B-5
B-6
B-7
B-8
B-9
B-10
B-ll
B-12
B-13
B-14
B-15
B-16
B-17
B-18
-------
LIST OF TABLES
Page
TABLE 1.
Geometry of Test Sections
. . . . . . . . . . . . .
25
TABLE 2.
Matrix of Test Conditions. .
. . . .
. . ...
51
TABLE 3.
Experimental Data for Test-Section Entrance
. . . .
58
TABLE 4.
Reduced Data for Test-Section Entrance
. . . . .
. . . .
58
TABLE 5. Experimental Data for Straight Test Section . 59
TABLE 6. Reduced Data for Straight Test Section . . . . . . . . . 59
TABLE 7. Experimental Data for 30-Degree-Bend Test Section . . . . 60
TABLE 8. Reduced Data for 30-Degree-Bend Test Section . . . . . . 60
TABLE 9.
Experimental Data for 90-Degree, 4-Inch-Radius-Bend
Test Section. . . . . . . . . . . . . . . .
61
TABLE 10. Reduced Data for 90-Degree, 4-Inch-Radius-Bend
Test Section. . . . . . . . . . . . . .
. . . . .
61
TABLE 11. Experimental Data for 90-Degree, O~Inch-Radius-Bend
Test Section. . . . . . . . . . . . . . . .
62
TABLE 12. Reduced Data for 90-Degree, O-Inch-Radius-Bend
Test Section. . . . . . . . . . .
. . . . .
62
TABLE 13. Experimental Data for Replication Test With 90-Degree,
4-Inch-Radius-Bend Test Section. . . . . . . . .
63
TABLE 14. Reduced Data for Replicati9n Test With 90-Degree,
4-Inch-Radius-Bend Test Section. . . . .
. . . .
63
TABLE 15. Experimental Data for Rough-Wall and Heated Test
Sections. . . . . . . . . . . . . .
. . . .
64
TABLE 16. Reduced Data for Rough-Wall and Heated Test Sections 64
TABLE 17. Fuel Mass-Balance Summary . . . . . . . . 65
TABLE 18. Wall-Temperature Data for Heated Test S ec tion . . . . . . 81
TABLE 19. Comparison of Alternative Atomization Systems 90
-------
PHASE II REPORT
on
A STUDY OF THE INFLUENCE OF FUEL ATOMIZATION,
VAPORIZATION, AND MIXING PROCESSES ON POLLUTANT
EMISSIONS FROM MOTOR-VEHICLE POWERPLANTS
by
D. A. Trayser, J. A. Gieseke, R. D. Fischer,
and F. A. Creswick
INTRODUCTION
This study was initiated in 1968 with an 8-month Phase I analytical
program to explore the incentives for achieving engine operation with leaner
mixtures as a means of reducing exhaust emissions, and to develop information
on potential means of extending the lean operating limit through improved
induction-system design concepts and improved analytical design approaches.
(1)*
A report on the Phase I program was issued on April 30, 1969.
Among the significant conclusions of the Phase I program were:
improving air-fuel distribution to extend the lean operating limit can lead
to worthwhile reductions in exhaust emissions, and the most promising approach
to achieving improved distribution is through use of fuel atomizing devices
that will produce droplet sizes approaching 10 or 20 microns under all operating
conditions and by design of the induction system for minimum impaction of fuel
droplets.
It was recommended at the conclusion of the study to extend the
investigation into a Phase II experimental program to demonstrate the technical
feasibility of improved induction-system design for reducing exhaust emissions~
It was also recommended that, as a part of this extended investigation, experi-
mental studies of fundamental induction phenomena be carried out.
In the course of conducting the Phase I study, it had become clear
that very little was known about droplet dynamics, wall fuel-film characteris-
tics, and vaporization as they apply to the induction system of an automotive
* References are listed at the end of the report.
-------
2
engine. Very little data are available on these phenomena, and even less has
been published on methods of measuring and/or evaluating these parameters.
Mathematical approaches to predicting droplet impaction and wall fuel-
film patterns are limited, for practical reasons, to idealized one-dimensional
steady air flow. Predicting fuel vaporization from fundamental data is
frustrated by the complexity of gasoline composition. Consequently, experimental
studies were needed to augment the simplified analytical studies that had been
conducted in Phase I, and to guide the selection of improved induction system
components.
OBJECTIVES
The objectives of this Phase II study were to obtain experimental
data on droplet impaction characteristics, fuel-film flow on manifold walls,
and fuel vaporization for a better understanding of induction-system phenomena;
and to evaluate the potential of improved fuel atomization, fuel vaporization,
and intake manifold design for improving air-fuel mixing and distribution.
SUMMARY OF. STUDY
The Phase II program was primarily experimental. The study included
three major tasks: development of instrumentation for measuring mixture proper-
ties, experimental studies of induction-system phenomena, and conception of
a prototype improved induction syste~. A fourth major task, construction and
demonstration of the prototype improved induction system, was originally part
of the planned program but had to be eliminated when the first two tasks
required more time and effort than was anticipated.
Development of Instrumentation
Instrumentation and procedures were investigated to measure the
amount of fuel in droplet and vapor form in the air-fuel mixture stream and
the amount of fuel on the manifold-passage walls.
A single-stage impactor
-------
3
probe was developed to separate the liquid and vapor while sampling from
the mixture. The sampled liquid and vapor streams were analyzed separately
for fuel concentration using a Flame Ionization Detector Hydrocarbon Analyzer.
An elaborate apparatus was set up to obtain accurate flow measurements of
the sampled gas quantities and to maintain all fuel in vapor form for the
analyzer. A previous method tried for fuel concentration measurements,
using gas-chromatography equipment, was unsuccessful.
After considerable data were accumulated using the impactor sample
probe it was found that, under certain conditions, the probe was intermittently
failing to fully separate the fuel vapor and droplets. At this p,oint a non-
separating sample probe was introduced and only the total entrained fuel was
sampled and measured.
A fuel skimmer system was developed to trap and measure the fuel on
the manifold-passage walls. With this system, fuel having deposited on the
walls of the manifold elbow or passage was pulled by vacuum through porous-
metal inserts in the walls at the passage outlet and subsequently condensed
and weighed.
Experimental Studies
Studies were conducted to measure the amount of fuel impacting on the
walls of various manifold geometries under different operating conditions. To
accomplish this, a laboratory flow system was developed to provide a simulation
of the fuel and air mixing and flow in an automotive induction system. Manifold
test sections of different bend radii, bend angle, and surface roughness were.
used in the experiments.
A spinning-disk-atomizer mixture generator was used
to generate fuel droplets of about 14 microns mean diameter for the laboratory
flow system.
Liquid-film transport on the manifold walls was observed visually
under the different geometry and operating conditions. The amount of fuel
flowing on the manifold walls was measured by the fuel skimmer system developed
in the first task of the program.
Fuel entrained in the mixture stream was
measured using the sampler probe and measurement system also developed in the
first task.
-------
4
Fuel vaporization from entrained droplets and from the manifold
walls was also studied to a limited extent in the laboratory flow system,
both by heating the air entering the apparatus and by means of a special
heated test section.
Improved Induction System Design Concept
Sketch layouts were prepared of a recommended improved fuel-
atomization system, a low-impaction intake manifold, and a fuel vaporization
chamber.
The fuel atomizer and intake manifold were sized to fit under the
hood of a standard size automobile.
The fuel vaporization chamber was con-
ceived as a laboratory experimental tool.
CONCLUSIONS
The results of this Phase II experimental study have led to the
following conclusions with respect to induction-system phenomena and directions
to be taken in future development efforts:
(1) Droplet impaction in an induction system can be
minimized by ultrafine atomization, minimum manifold-
passage turning angle, long bend radii, and low air
velocity. However, it does not appear possible to avoid
appreciable droplet impaction, even with mean droplet
sizes as low as 14 microns, because of deposition by
flow-induced air turbulence.
Twenty to fifty percent
droplet impaction, depending upon air velocity, was
observed in a I-foot long straight test section with
l4-micron droplets.
(2) Accordingly, while improved fuel atomization and the design
of manifold passages for low droplet impaction can probably
improve mixture distribution, these design features alone
are not sufficient to avoid an appreciable fuel-film
accumulation on the passage walls, and the attendant fuel
"hangup" and possibility of maldistribution. Therefore, it
is concluded that fuel vaporization should also be employed
-------
5
to some extent in the design of advanced carbureted induc-
tion systems.
(3) Fortunately, u1trafine fuel atomization also promotes rapid
fuel-droplet evaporation.
It appears that 14-micron fuel
droplets, as used principally in this study, may approach
complete vaporization in residence times typical of low-
speed engine operation with current induction-system design.
Also, it appears that droplet vaporization may be appreciable
even at in-manifold residence times consistent with moderately
high-speed engine operation.
Accordingly, the use of preheated
intake air, preferably above the dew point of the mixture, is
highly desirable. (The increase in air density resulting
from the cooling effect of fuel vaporization more than off-
sets the displacement of air by fuel vapor and should result
in an increase in engine volumetric efficiency.)
(4) Tests conducted with a test section heated to 190 F resulted
in effectively alleviating most of the wall fuel-film that
would have accumulated under conditions when droplet impaction
would otherwise have been substantial. This is'probably due
to the combined effects of vaporization directly off the test-
section wall and increased droplet-vaporization resulting from
increased air temperature. Accordingly, it is concluded that
moderate, uniform manifold heat is adequate to virtually
eliminate the unavoidable wall fuel-film in an improved-
atomization, low-impaction system.
(5) The most promising approach to the design of a carburet ion-
type induction system appears to be through the use of the
following features:
. A fuel atomizer capable of producing droplet sizes
in the range of 20 microns and smaller.
. The use of preheated intake air (100 F or higher).
. Intake manifold design for minimum turning angle,
long bend radii, and moderate air velocity.
. Good mixing and long residence time in the mixture
generator, ahead of the manifold.
-------
6
. Moderate» uniform manifold heat (in the range of 190 F).
(6) The design approach preferred by the project staff
incorporates the following design features:
. An air-atomizing fuel nozzle operating on
the ultrasonic Hartmann-whistle principle.
. A mixing section similar in construction to a
can-type combustor.
. An unconventional throttle valve: either a plug
valve between the mixture generator and intake
. manifold or a butterfly valve upstream of the
mixture generator.
. A single-plane intake manifold with a separate
branch for each cylinder having equally spaced»
equal-area, radial entrances.
(7) Experimental techniques for sampling fuel/air mixture
properties need further development.
Our understanding of
droplet impaction phenomena is far from complete; however,
further knowledge in this area is not of crucial importance
at present» since it is evident that droplet impaction cannot
be avoided completely under any circumstances. Further
experimental information on fuel-droplet vaporization could
be of considerable help at present in advanced induction-
system design.
RECOMMENDATIONS
(1) Experimental studies of fuel-droplet transport and vaporization
should be continued in support of advanced induction-system development. Fuel-
droplet vaporization should receive primary emphasis at this time.
(2) Design and development of advanced induction-system concepts
based on present knowledge should be carried out expeditiously.
Detail design,
fabrication, and experimental evaluation of the preferred system concept
presented in this report is recommended.
-------
7
EXPERIMENTAL WORK
The objective of the experimental program was to determine, by
actual measurements, the effects of droplet size, manifold geometry, and
operating conditions on droplet impaction, wall-film characteristics and
fuel vaporization characteristics.
The approach taken to accomplish this
was to construct an air- and fuel-flow system to simulate the mixing and
flow of air-fuel mixtures through manifold passages.
A system was developed to accomplish these objectives. However,
actually obtaining meaningful measurements of mixture quality turned out to
be far more difficult than anticipated. Mapy problems were encountered,
both in developing the air- and fuel-flow system and in developing measure-
ment techniques.
As a result, not all of the objectives of the program were
fully met. However, it is believed that the experimental results that were
obtained represent new and useful insights into induction system phenomenon.
The topics covered in this section of the report include: the
laboratory flow system manifold test-section geometries, fluids used in
the experiments, mixture quality measurement, the experimental studies per-
formed, and the results of these studies.
The Laboratory Flow System
Performance Criteria
The aim of the laboratory flow system was to simulate air and fuel
flow characteristics as they occur in real induction systems, but under con-
ditions that could be controlled,measured, and varied. Specifically, it was
desired to simulate the flow through a single branch of the manifold on a
300 eIn engine from idle to maximum speed for different engine load conditions,
over an air/fuel ratio range from about 12:1 to 20:1, and using fuel droplets
from about 20 ~ up to at least 100 ~ in diameter. Additionally, it was desired
to control the temperature of the air entering the system over a range from
80 F to 150 F.
-------
,-~.
8
Descriptions of Components
Figure 1 is a schematic of the laboratory flow system showing the
major components and subsystems.
The major components are: the air flow
system, the fuel system, the mixture generator, the air preheater, and the
discharge burner.
These will be described in the following paragraphs.
Shown
also in Figure 1 are the test section and the sample-probe section.
described in other sections of the report.
These are
Air Flow System. The air flow system consists of a flow-inlet section,
a flow-control section, and a high-vacuum p~mp.
Air flow through the laboratory system is filtered and metered at the
flow-inlet section. Four dry-type automotive air filters were clamped together
to provide effective filtration with low pressure loss. At the maximum system
air flow of 93 scfm, this pressure loss is 5.3 in. H20.
Average system air-flow rate is measured with an ASME square-edged
orifice meter in a 1-1/2-inch pipe with 1 D and 1/2 D taps. An air-flow range of
3 to 100 scfm can be measured with this meter with an accuracy of to.75 percent
using one of three orifices having orifice diameters of 0.400, 0.860, and 1.239
inch. With the largest orifice, the differential pressure at maximum system air
flow is 15.3 in. H20.
Computed flow vs pressure differential data for the orifices used in
the experimental program are presented in Appendix B, Figure B-1. These data
were calculated using equations from Fluid Meters(2) for standard ASME square-
edged orifice meters.
Vacuum in the test section is controlled by a gate valve located down-
stream of the air-flow orifice meter.
Higher vacuum can be obtained by closing
this valve.
With the valve wide open, minimum vacuum in the test section is
fixed by the flow losses from the air-filter inlet to the test-section inlet.
At the maximum system air flow, this minimum vacuum in the test section due to
flow losses is 2.7 in. Hg.
-------
.~ Test section
vacuum control
valve
J'
\
)
, ..-
\,flow ---.'
Inlet
Air flow
orifice meter
Inlet air
filters
FIGURE I.
Air heater control
transformer
Air preheater
Mixture generator
Shop air-
.Air motor
control valve
Transition, round-to-square
Test section
Vapor-
Vapor/droplet sampler I
Liquid'
Sampler section
----- Oil injection
Motorized boll valve
Fuel pump
Fuel flow orifice meter (high flows)
Solenoid shutoff valve
\0
Burner
High-vacuum
pump
SCHEMATIC OF LABORATORY FLOW SYSTEM FOR INDUCTION
SYSTEM FUNDAMENTAL STUDIES
-------
10
A ball valve driven with an electric motor is used to produce pulsating
flow in the test section. A standard 2-in. ball valve with "Teflon" seats was
modified for use at driven speeds up to 1400 rpm. A central shaft was added to
support the ball, which is normally supported only by the "Teflon" seats. The
seats were lightly loaded against the ball with wave springs. A small amount of
oil is injected into the air flow system immediately upstream of the ball valve
to lubricate the ball.
The performance of the ball valve was evaluated in the laboratory flow
system with the flow and vacuum control valves set for maximum flow. The ball
valve was operated at several different speeds and the flow velocity was measured
at the test section inlet with a Thermal Systems hot-wire anemometer.
flow rates were measured with the flow system orifice.
Average
Figure 2 shows oscilloscope traces of the flow velocity-time data from
the hot-wire anemometer. Average flow rates for the six ball-valve speeds are
as follows: 62.5 scfm at 400 rpm, 58 scfm at 760 rpm, 66.5 scfm at 800 rpm,
69.5 scfm at 1000 rpm, 68 scfm at 1200 rpm, and 64.5 scfm at 1400 rpm. The
average flow increased with increasing speed up to 1000 rpm. Beyond that speed
the average flow decreased slightly. An operating speed of 1000 rpm was selected
for the test program as yielding the maximum average air flow with an acceptable
velocity-time curve shape.
Figure 3 shows that the time variation of the pulsating flow produced
by the ball valve is expected to simulate something between that existing in a
manifold runner and that existing in a manifold branch. At 1000 rpm the ball
valve will produce a pulsating frequ~ncy equivalent to that in a manifold branch
at 4000 rpm engine speed or in a manifold runner at 2000 rpm engine speed.
A 30-gallon surge tank is used to dampen the flow surges caused by the
vacuum pump. At the maximum system flow, the amplitude of the pressure pulses
in the surge tank was calculated to be about 1/2-in. H20.
Air flow in the laboratory apparatus is controlled with a flow-control
valve mounted between the vacuum pump and the surge tank as shown on Figure 1.
Maximum system air flow is obtained with this valve wide open and is a function
of the capacity of the vacuum pump and the system pressure losses. Flow can be
reduced by partially closing the valve and thereby reducing the density of air
-------
) -
250
~ 200
U)
........
-
-
>: .50
-
u
o 100
4)
>
~ 50
o
~
11
o
.
.
~
0.02 see
Ti m e ---..
250
u 200
4)
.!!!
.:: 150
:>:
-
u
.2 100
~
~
o
u...
o
"--v---'
0.01 see
Time ---.
FIGURE 2. OSCILLOSCOPE TRACES OF PULSATING FLOW
-------
12
Idealized
f.low in
manifold
runner
Idealized
flow in
manif.old
bronch
Pulsating flow
simulation
generated with
rotating valve
FIGURE 3.
TIME VARIATION OF PULSATING FLOW
-------
13
at the inlet to the pump.
System air flow is directly proportional to air
density at the vacuum pump inlet because the pump is a constant-volumetric-
displ.acement machine.
Air is pumped through the laboratory system by a rotary-vane, high-
vacuum pump which was available in the lab. The pump is equipped with a closed-
cycle, cooling-oil system for high vacuum operation. For use in the laboratory
flow system, the oil that is injected into the pump inlet is not collected and
returned to the pump but is burned with the fuel vapor. Proper vane lubrication
can be achieved with a very low oil flow.
The pump volume flow is relatively constant at 115 cfm referenced to
the inlet vacuum condition.
A vacuum of 20 in. Hg or lower can be produced at
the test section.
The maximum system air flow with minimum system restriction
is 93 scfm or 7.1 lb/min.
Figure 4 shows the relationship between air flow and manifold vacuum
for the laboratory flow system operating ~t both steady and pulsating flow
conditions. To develop these curves, the flow control valve was fully opened
and the vacuum control valve was adjusted at various settings from nearly
closed to fully opened.
The vacuum was measured at the manifold test section.
The representative carburetor flow curve included on Figure 4 depicts
the flow through one barrel of a two-barrel carburetor fitted on a 283 CID V-8
engine.
The air flow capacity of the laboratory flow system approximately
matches this representative carburetor flow curve with pulsating flow, and
greatly exceeds it at steady flow.
Other types of air pumps were considered in the design of the laboratory
flow system but none were considered as suitable as the present arrangement with
a steady-flow vacuum pump and motorized ball valve.
One alternative arrangement
considered was a motored automobile engine equipped with a modified manifold.
This setup was judged to be a relatively expensive installation even if set up
with an inexpensive, used six-cylinder engine because of the setup time and the
long-term use of a motoring dynamometer facility.
The engine could be motored
with a 10 to 15 hp electric motor but the speed range would then be limited.
Manifolds would have to be fabricated to connect the intake ports of the active
cylinders to the test section and the exhaust ports to the fuel-vapor burner.
-------
14
Another alternative arrangement of air pump considered was a large,
reciprocating vacuum pump. Rotational speed of most reciprocating vacuum pumps
is limited to about 1000 rpm, which would provide simulation of the flow in a
manifold branch at 2000 rpm engine speed, but would provide simulation of the
pulse frequency in a manifold runner at only 1000 rpm engine speed.
In addi-
tion, single cylinder compressors with adequate capacity were judged to be too
expensive.
24
20
\ '\
\\
16 \\
CI ~
I
III \
Q)
~
u
c 12 '\
E \\
::J
::J \\\
u
0
> \\ \
8
Representative ~\ \
carburetor flow Y
4 \, \
Pulsating flow \\ \
\ \
,
\
\
00 20 40 60 80 100
. Flow, scfm
FIGURE 4.
AIR FLOW CAPACITY OF LABORATORY FLOW SYSTEM
-------
15
An arrangement with two inexpensive two-cylinder compressors coupled
together with a timing belt was also considered.
Compressor intakes could be
connected together to provide simulation of the wave forms of the pulsating flow
in either a,manifold branch or runner. Operation at 1500 rpm was felt to be
feasible by increasing the size of the intake valves.
Further consideration
of this two-compressor arrangement was dropped when the large rotary-vane
vacuum pump became available.
Fuel System. Fuel is supplied to the atomizer in the laboratory
system with a centrifugal transfer pump which has a stable flow range of 0 to
240 gal/hr. Pressure rise at zero flow is 13 ft of fluid being pumped. Seal
leakage between motor and impeller is eliminated with a magnetic coupling.
Fuel is pumped from a 55-gallon container, installed outside the
building, into a heat sink to bring the fuel to room temperature. Following
the heat sink, which consists of a coil of 1/4-in. copper tubing, the fuel is
filtered in an automotive fuel filter. Fuel flow rate is controlled with a
precision flow control valve following the filter, and is measured with a rotameter.
A fuel rate of about 4.6 gal/hr is required to provide a 15:1 air-fuel ratio at
the test section at maximum system air flow. Following the meter, the fuel flows
through a solenoid shut-off valve and into the atomizer. This valve is used to
prevent the fuel from draining back to the supply tank when the transfer pump
is shut off.
A calibration curve on the rotameter used for fuel flow measurement is
presented in Appendix B, Figure B-2.
Stoddard Solvent.
This calibration curve was obtained using
Mixture Generator.
Figure 5 is a layout drawing of the mixture generator
showing locations and orientations of its components. The housing consists of a
cylindrical center section with conical expansion and contraction sections upstream
and downstream. The cylindrical portion contains a spinning-disk atomizer with
supporting members and a fuel feed system. A collector is provided at the outlet
to remove fuel from the flow system which has impacted on the wall of the mixture
generator.
The overall length of the system is 24 inches.
All joints are leak-
tight with neoprene gaskets or O-ring glands.
-------
8
H
G
F
E
D
c
B
A
.
7
e
II
4
I l
! '
. I
"TI
G)
C
::0
fTl
~
r
~
o
c
-;
o
::0
~
Z
GJ
o
"TI
s:
X
-;
C
::0
fTl
GJ
fTl
Z
fTl
::0
~
o
::0
~
~
~
~r
-0
A<'t:)~ ~ A¥l.LT-
7
e
.
~.
~,-~
~hV"
8
e
-~ """'... n/~r .-""'"
'~r~~
3
\
---\
~
2
O/~-3"I'.NcS'"
a / ~£r ~£C77t:VV'
~ / GV./TZ.cr.$£Cr7QN
,I(/J. J:< ere ~~ /2~E
...s z It:; TZ/6£-~£
& ..., / 72/.8£ A'M~~A:'
7 ~-/ J!;.;;;;,v~---
~ - .~; o.D..r--O~.5JI'Y. r. e;nr:s.:J W6£':io!--
'? - - I c.Cit.r.~.5Nr.~.:J",7V4£-~-'<:6.-
OO..:JGILJ I' /N...IECno"';'-T£E
// /,.8 / h'V../,cC?'7a-v NOZZLE
lo!'j04:UZ!8'/ ~~£r o.<$C
/..$ -1- ~.TEE .7 n/4E .r.T~~ ...,~r
5~~':::0~;..," G
,/. ,- -~:Lfo.QA'~~~-=~=~
./ ,- - ",=...~.:!/:-_L.& .so:: -:":'.0. e:;,,- J~
- .-.1$ "10-<&''''A'~H.Q ..5ETJ.c.
,19: _~~8£--
- - ~ ~4 ~~r~~...~o~: ~~~:~~~ -
~/ - - I I #'0 -"4'0 Rh#:" -vp,nu:
H
----
F
~ ~,
~;
t ,.1.........,--
-"'.,a/uJ".~r
~ ~
jJ
E
\
......
0'
-r=TLAS.r.:wco; T?/~r
D
c
"' B
MT'ftI.U ~.-nnn8
---
~.-
.;::-:..---- .-- - -"'~"""'NN//\0t::7 .oA:iC
: I ~z~
- ~t:t.rT4-/ i'OC z E
A
2
-------
17
The center cylindrical section is 9 inches in diameter and formed
from "Lucite". Fuel which does deposit on the walls is removed at the exit
from the downstream conical section by
a "skimmer".
This is simply an annular
space into which the fuel is collected externally in a small chamber where it
can be measured either by weight or volume. Pressure equilization between
this chamber and the mixture chamber allows the fuel to flow freely.
Figure 6 is a detail sketch of the fuel skimmer system. Whi+e
setting up equilibrium conditions for a run, Valve 1 is closed and Valve 2 is
open so that fuel collecting in the skimmer is continuously purged out. When
making a run at steady state conditions, Valve 2 is closed and Valve I is open.
The vacuum draws fuel into the graduated cylinder from the skimmer. The time
is recorded to reach a predetermined volume. The rate at which fuel is drawn
from the skimmer chamber by the vacuum is adjusted carefully so that only
liquid fuel and not vapor is pulled from the mixture chamber. This is accom-
plished by maintaining the skimmer fuel flow just below the point where bubbles
begin to appear.
The major component in the mixture generator is the spinning-disk
atomizer. The aluminum disk itself was machined in the shape of a low-height
inverted cone with a thickness of 15 mils at the outside edge and an integral
shaft on the underside. The disk is driven by an air turbine at speeds up to
85,000 rpm.
The head or top portion of a commercially available air turbine*
was replaced with a specially designed section which traps air leakage from
the turbine and prevents its escape ~nto the mixture generator. A vacuum pump
is used to maintain a negative turbine-exhaust pressure to increase the maximuffi-
speed capability of the turbine and to make it possible to adjust the pressure
in the turbine head so that it equals the pressure in the mixture generator.
Pressure taps in the turbine head and the wall of the mixture generator are
connected to a manometer so that the pressure difference can be adjusted to zero.
The motive air for the turbine is controlled up to 90 psig and enters
at the bottom of the turbine through one of the supporting tubes. A second
sYmmetrical supporting tube at the bottom serves as a line to the manometer for
the head-pressure equalization scheme.
tubes at the upper end of the turbine.
The motive air exits through supporting
* ARO Corporation Model No. 7980.
-------
18
Mixture generator
lower section
Collector annulus
Tal'-,
vacuu~
v
Mixture
flaw
3/4-in. ID graduated cylinder
To ~
vacuum
FIGURE 6.
DETAIL OF MIXTURE GENERATOR FUEL SKIMMER SYSTEM
-------
19
The rotational apeed of the turbine is monitored by a Strobotac
connected to an electronic counter.
The fuel enters through a tube above the spinning-disk and passes down
onto the .center of the disk through a hypodermic needle. It is tmportant that
this needle be located at the center of the disk; otherwise uneven and slugging'
atomization results.
The range of fuel-feed rates possible with this system is up, to
50 lb/hr. An upper ltmit of 25 lb/hr is required to provide an air-fuel
ratio of 20:1 for the air-flow rates planned for the experimental program.
The .pinning-d1sk atomizer proved to be unacceptably noisy when
operated at the high apeeds which were nece~sary to obtain small droplet sizes.
After .pending .ome time attempting to pinpoint the cause of the high noise
level and attempting to iaolate the atomizer from the mixture-generator housing
c08ponents, we concluded that the only solution was to enclose the entire system.
To this end a "aoundproof" box was built around the mixture generator including
the coaical .etal tranaition sections before and after the atomizer housing.
Thi. box conai.ted of an outer shell of 3/4-inch plywood, a layer of 0.065-inch
l..d .heet, and 2 inches of foam plastic. Th~ box was made in two halves hinged
to..ther 80 that acceas to the mixture generator for service and adjustments
would be facilitated. A window on one side permits observation of the spray
patterna and 8Onitoring of the disk speed when in operation.
Other problems with the spinning~disk atomizer included an air-motor
bearina failure and air leakage at the top bearing. The failed bearing proved
to be defective and was replaced. T~e air-leakage problem was solved by attach-
ina a vacuUII p..p to the air motor exhaust and to a "trapping" chamber built
oato the air motor at the top bearing. The vacuum in this trapping chamber was
aaintained equal to the vacuum in the atomizer chamber by means of a control
valve, 80 that no leakage of air or air-fuel mixture would occur between the
two .paces. .
An attempt was made to operate the air motor and disk upside down so
that tube. and supports in the mixture flow path could be eliminated. A uniform
.pray pattern could not be achieved under these conditions.
-------
20
A large percentage of the fuel atomized by the spinning disk deposited
on the mixture generator walls before it could enter' the' test section. This was
particularly true at low flow rates. The diameter" of the mixture generator"'
housing was originally selected after observing the atomization pattern in
quiescent air. However, the droplet trajectories turned out to be quite differ-
ent when the atomizer was operated in the mixture generator.
The primary causes
of the high rate of impaction appeared to be the relatively low axial ,air-flow
velocity in the chamber, a radial air flow created by the spinning disk, and a
swirl pattern also created by the spinning disk.
A means was finally devised to eliminate or minimize the above-
mentioned phenomena.
Modifications were mqde to the mixture generator and the
results appeared favorable.
Figure 7 is a sketch of the mixture generator center section showing
i~ ".
the modifications.
A 6-1/4-inch diameter baffle was installed in 'the center of
the chamber in two sections, one above the disk and the other below 'the disk.
A I-inch spacing was provided between the two sections for the spinning disk
and fuel spray.
Each section was split in an axial plane so it could be installed
without disturbing the air motor and tubing. Both sections were 'closed'and
sealed at both ends. Six equally spaced vanes were longitudinally mounted in
the 1-3/8-inch annular flow passage between the baffle and chamber wall.
The narrow space between the baffle sections provided relatively
quiescent air for the fuel spray to pass through which reduced their trajectory
length. The smaller flow area provided by the annulus passage increases the air
and mixture velocity considerably, a factor which would also reduce the amount of
vaporization occurring in the mixture generator. The longitudinal vanes prevented
a swirl pattern from developing, although they undoubtedly 'become impaction sur-
faces in themselves. Loss of fuel droplets to the mixture generator walls was
considerably reduced by the baffle and guide vanes, although the size of the
droplets entering the manifold test sections appeared to be unchanged;
Air Preheater.
Air temperature at the test section is controlled by
heating the air to a constant value in an electrical resistance heater.
The
heater is a 4-foot length of 1-1/2-in. diameter pipe which is wrapped with two
-------
21
Vanes (6)
Baffle
Baffle
FIGURE 7. MODIFICATION TO MIXTURE GENERATOR
TO. REDUCE IMPACTION
-------
22
768-watt heating tapes and insulated with 1-l/2-inch thick high-temperature
insulation. Maximum heater-tape temperature is 900 F. Power input to the heater
is controlled with a Variac autotransformer, which varies the voltage input.
Temperature of the room air can be increased 59 F at 53 scfm and 41 F
at 90 scfm.
Heater effectiveness, expressed as heat transferred to the air
divided by input power, is 73 percent at 53 scfm and 86 percent at 90 scfm.
Discharge Burner.
Fuel and oil vapors are disposed of in a burner
connected to the vacuum pump outlet.
The burner is set up with a continuous
natural-gas pilot flame and has a weighted poppet valve at the inlet to prevent
flash-back.
A flame arrestor is installed between the burner and the vacuum pump
as an additional precaution against explosions.
The flame arrestor has an
aluminum matrix which quenches the flame by cooling.
Figure 8 is a sketch of the burner showing the anti-flashback poppet
valve.
Manifold Test Section Geometries
Requirements
One objective of the experimental work was to obtain verification of
the analytical predictions of droplet impaction, fuel-film, and vaporization
characteristics made during the Phase I study. Consequently, the manifold
geometries selected for this study were based on the geometries described in
the Phase I Report of April 30, 1969. (1)
It was also required that the pas~age shape be similar to actual
intake manifold passages, and that the cross-sectional area be selected in
conjunction with the air flow system capacity to yield a maximum flow velocity
equivalent to that which occurs in real manifolds.
Inside surfaces of real manifolds are rough from the casting process.
Consideration was given to attempting to duplicate this roughness in the experi-
mental manifold test sections. It was believed that surface condition has very
little influence on the droplet impaction phenomenon.
Therefore, it was decided
-------
23
Gas valve
--~
/
Burner can
Poppet valve
Teflon guide
bea ri ng
Mixture flow A
from y
flame arrester
Teflon guide seal
0.90 piano wire
FIGURE 8.
DETAIL OF DISCHARGE BURNER
-------
24
to select the material for the majority of the test sections for other
considerations such as ease of construction and transparency for observing
the wall fuel-film flow patterns. However, surface roughness could playa
significant role in fuel-film transport and in reentrainment, thus, one test
section was provided with rough surfaces.
Description of Manifold Test Sections
Test sections for the impaction studies were fabricated from 1/4-inch
thick, clear acrylic-plastic sheets glued together to form a square duct. Side
walls were machined with the circular-bend.outline and then the top and bottom
walls were heated and bent to conform to the side wall radii.
With this con-
struction, elbows with constant cross sectional area with the desired bend radii
were easily fabricated.
These test sections with square cross section provide
a reasonable simulation of rectangular intake-manifold passages used on many
engines. The transparent walls facilitated visual observations of flow phenomena.
Table 1 gives pertinent geometry of the available test sections. The
1-1/4-inch width was judged to be about minimum size compatible with practical
vapor/droplet sampling probes.
The length of each test section was fixed at
1.0 ft to approximate the length of a single manifold passage in a V-8 engine
intake manifold. All test sections were made the same length so that droplet
impaction measurements would reflect only the influence of bend angle, radius
and wall surface texture.
As noted in Table 1, the walls of one of the test sections were
roughened to simulate a sand casting.
This was accomplished by brushing the
acrylic plastic with
section duplicated a
could be made of the
a steel brush soaked in acetone solv.ent. The rough-surface
test section with smooth surfaces so that a direct comparison
effects of surface condition.
Drawings of the clear plastic manifold test sections used in the exper-
imental program are shown in Figures 9A to 9C.
The aluminum instrumented test section listed in Table I is also a
duplicate of one of the clear-plastic test sections.
This section, when mounted
in the flow system, is wrapped with electrical heating tape so that all surfaces
-------
25
can be heated to observe the effects of wall fuel-film vaporization.
Ten
thermocouples are mounted at strategic locations on the test-section surfaces
to measure the wall temperatures.
thermocouples.
Figure 10 shows the locations of these
TABLE 1.
GEOMETRY OF TEST SECTIONS
Cross Duct Outer Bend Bend
Section Width, Radii, Angle, Surface
Shape in. in. deg. Texture Comments
Square 1-1/4 Infini te 0 Smooth Straight duct for
reference
Square 1-1/4 0 90 Smooth
Square 1-1/4 4 90 Smooth
Square 1-1/4 4-7/8 30 Smooth
Square 1-1/4 .4 90 Rough
Square 1-1/4 4 90 Smooth Aluminum instrumented
test section for
manifold heating
tests
-------
'~
-,~
"-,,.~ ~>, '.
--<~~ -5. o:ri J
--. J:- . . 500 ~ i
'':::::: .. <'.., -t-=-'-'=c=., '='= -- -- - -~ ~
<,~---~- -- J - - -- -- - - - -- . - - - - _:::
"- -----' .~-_.- ~------_._- --------
30' ---~-::... .--- --., 'u- u_-- ~
-j
!
FIGURE 9A. 30-DEGREE, 4-7/8-INCH RADIUS MANIFOLD TEST SECTION
4.50
--""'5-40/YC-C.8 )( .3/ DEE?
<5 HOC.E5 LOCh'TE,o r-/FO/W'
?h'£'T 00// 6' Eh'Ch' EN,o
U. .~~_l
~ .:,
~ ..' $ '-./~TY~
.~ .
N
0\
~ - .e5 T,Yr.<-
i
i
,
f--/. 7.5---i
-------
27
i I
I :
I
I! '
I ,
7./C
i!
I i I
I: 'L
l
/...50
.L
C:;;..5" 8
FIGURE 98. 9O-DEGREE, O-INCH RADIUS MANIFOLD TEST SECTION
\#.5-40/VC-~8 X.-3/occr"
~ HOLC':' t.OCHT£O ;:;;roM'
~RT 00// e E:HCNEND
--L
'$
$
e
.IC TYP
..1--.
$
$
$ r3)
I J L.~.5 TYr.'"
~/.7S ~
-------
~ : :;
~:- !~ ~
!
/.70
\
I: 7-
I I
\ \, / :
\ \ , ' I
\ .~\\\ 4./c.€"./ I'
\ '
\ // .., eGzc,e-l.
\ .~
..5.00
~I
~#..5-40/VC_C.6" A' .-3'/ oee?
6' .-7'OLES LOCh'TEO ~~Oh?'
P/?RT 00/ / ~ c/?CH eND
I
N
00
4;--
-------
r?
5
10~
22-
2
29
Mixture
flow
D
Note: AI I thermocouple junctions
cemented to walls of test
section in IfIG-in. deep
cavities at midpoint
4
45°
67~
2
90°
No. I thermocouple located
. on backside opposite
No.8 the rmocouple
FIGURE 10.
SKETCH OF ALUMINUM INSTRUMENTED TEST SECTION
. SHOWING THERMOCOUPLE LOCATIONS
-------
30
Discussion of Fluids
The impaction of gasoline on the walls of a manifold can be pre-
dicted by the use of dimensionless equations which have been shown to provide
proper correlations. Except for density, the physical properties of the fluid
do not apply in the correlations, Consequently, if densities are known and
are reasonably close numerically, an alternate fluid can be successfully used
for droplet impaction studies. For instance, the density of Stoddard Solvent
3 3
is 48.8 lb/ft compared to 44.9 lb/ft for gasoline, thus, any correction for
density would be small and probably not significant with respect to practical
considerations.
A dimensional analysis indicates that three dimensionless groups
control droplet impaction. (3) These are
. Impaction parameter
K = (27Tn)1/2 [ PdUg J
l8~ D
g m
1/2
Dd
. Parameter characterizing
non-Stokes law behavior
2 D
cI> = 9P g u g m
Pd~g
. Geometry of bend
~/Dm
Where:
n = number or fraction of 360-degree turns in bend
Pd = density of the droplet
u = velocity of the gas
g
~ = dynamic viscosity of the gas
g
D = manifold passage diameter
m
Dd = droplet diameter
Pg = density of the gas
Rb = bend radius.
-------
31
The case for an ideal gas has been analyzed by Ranz(3) and the results have
been presented as a dimensionless correlation of impaction efficiency in two-
dimensional ducts as a function of the impaction parameter, K, for various
2 2
values of the combined parameter, 4> / (27rn) (Rb/Dm) .
Correlations for droplet impaction efficiencies using the dimension-
less groups described above serve to predict deposition on the walls.
Since
this is the objective of the droplet impaction studies, the correlating tech-
niques should prove to be satisfactory. There are, however, other phenomena
occurring which cannot be modeled in such a direct manner.
The more difficult problems are concerned with reentrainment of
drops from the tube walls, the flow in a liquid film along the walls, and
vaporization of the fuel. These, of course, could be studied directly with
a gasoline-air system. However, the important physical properties of gasoline
can be identified and matched with a substitute fluid, or alternatively,
differences in fluid properties can be accounted for by proper analysis.
Absolute vaporization rates should be studied directly with gasoline
because the multi-component nature of gasoline makes this process difficult to
analyze.
The multi-component composition gives rise to vaporization rates
which vary as a function of amount vaporized since the vapor pressure changes
with the amount vaporized.
However, knowledge of vapor pressures for various
fluids will allow compensations to be made in vaporization rates if fluids
other than gasoline are used.
In the additional areas of concern, film transport and reentrain-
ment will be the same for an alternate liquid if the liquid viscosity and
surface tension are similar to those for gasoline. Figure 11 shows a com-
parison between the viscosities of gasoline and n-heptane, one of the hydro-
carbon compounds in gasoline. It is evident that in the temperature range of
interest (60 - 80 F), the two fluids have viscosities that are nearly identical.
111
Viscosity data for Stoddard Solvent were not available; however, its wide-
spread use as a test fluid in carburetor development work would suggest that
it has suitable properties.
A comparison between the surface tensions for
Stoddard Solvent, n-heptane, and gasoline is not easily made, again because
of a lack of good data.
However, most hydrocarbons of the types predominating
in gasoline and Stoddard Solvent have very similar surface-tension properties.
-------
32
Q)
.!!!
o
a.
1.0
~1+: ~ .:: I I I!,! i I; I II l-W I II It. I i II I I 1111 I I ! i I i I I: I ! ! . i I i III I
O.gr"..:.;' ,iil "!' ii, I" 1\11 I'"~ '" II!I 'iI'll! ',; !Iiil \,\ II I ' -Ii i'" "I I';! I !I', II!! ,'"III!,
';";':1 ~4!1:iili':i +i+i--4W-illJ .llL ,i,il dil IIi! ,I 'I' )11 1:11 ,!Iil i",' 44il,i', ili!, :'U" ill! !,i! l.!,li ,
........ ',! : I' .-mttr !, II II j: '!i' II I Till II!! ;,! I ill! i! I ! iil! I, ' : I 11 I! ! i: I I : I ,'! I II ! I j TIll! \ ! i : I ! ! Ii i; I! I
0,8
..-i: ~ ,.. ,'I j !: i Iii! i WI !',! Ii:! ! 'i I!! 1 It!: llil ~.4~ i ill ill! i 1 i I I i I! JlW Ii!: i i ! i UJ i llu :j i i :':: Ii j i 'i,
-;:: -,', '< "'r'T 0 'FTT !!!:~~1 Iii! l:il j!ii' :1lTTrF [!!rHi!T ill: 11Tr!1TI iii' rtt mT ITIT TTn liit TttrtH m
0.7
~ ;.:,' i ! i : 1 i i ! i i i ! ! ! i i i : i I L' 't! ! i !! ! : i; i I !! !! i : ;! i! !!!; i I !! 1'!: i :! '! : : i \: !!!' " Iii i !;! j
_.-, ~ ..'-.~c__..... ~~ ..... '"rOo' . .-" -.- .'1" ".,. .-.- ...- ,.w., ... ,~. 11".. ,..,.4 t... :-rr." TT"
'" """ :i:1 oj: s;,U.lllilliiii ~ :!i! i:ii Ii!! JJi 11!1 i!! iiii ilii iii' ii!! 1:1i ii!! !!, !iii ;:!i il!: i;ii
0.6 j~U ~,LU ill: i::: * ~!it :j~l liU WE !Hji \~;!,4~~ LlUW'j:J Ull W: ~ll EU i-LU Iri1ij ml LUUij WI
0.5::' ti ::i! ii!! ~i;i ;:\i iL:'-~ Ii:: :1;: iJi! Iii ::11 Iii! ilil !I , ;Ii; :ili !!ii !\I, I!ij Ii:! iii: :Ii: ii!! Iii!
i : : : . . I' :: i J' ",. ;;:: ;!!: ; i :" ~ .......:. I: ~! il' ,III: ;: I,' ~ Ii :: i II: !; I" L' I : I : 'I i " : " ~; I!:: :::: I 'I' : i :;:: i::: :::: :: I I
".~ :::: ~; ::1,' :1il ,I:; :!I; ,", :::'. ..........:.:.~~ ,:,! :i : :i ! :: Ii! ~I :,:, I,:: il:: ::1; : l.i I::; j': ,::: ,:I!
:~" - " , :1:; :Ti"i:"; :I': :'1~': Ii;: :8::: ~ ':;, ~ :':11 iTT!iTf'l !P,"j 'ilT 1T:: :Tii I'!: fin ~11:;TT71; ':i-:: tii; ~t:
, .., "I, , " 'I': 011: ,,', , I: ' II', "" , .. h....' 1 01" "";,,.!, "I, "I 'I'! III ,I" "II "" ,,,, ": i II! I 'I'! III' .. I '
0.4;:!:~: :;~llliJ,L,!ii i~ ~~illi !:~ iil.,~d{:ilitt:~0 ~,ll; ~~ llU illi ill! ;ll\ ii!! !Iii 11i1, $:';: ;;1\
I::: :,:;; mi :!:! :1111, :111'1 :!'I !1111'1 111111i DI: :\'1': :'I!! %,,:,111111' iilT 'l'ii;,'~IMili, :1!: i!i !I:'I~-I~~~P~~~I~"'I'" 'I I I,
.~.. ..1\ .,!! 'I' . "I I I I. '1111 II ! 1 '1 t 1'1 !Ill "I 'I! 11"~' .L., I t' ", "." 1III!tI.il:.:1
0.3 ,:;!' ;i;: I!!: :1 I !!I~; i\'I\I\i !lli iii: 1;11 ill I II ! 1!lii! li~ill!I'liil II Ii ill'; :!i! >.J?:Gas;o,i~~:" :;:~!I'II WI: Ii
:::i :l!' 1:1: illl'\I: 11"11 Ijl', '~liil~1 ml1il j11'I' i~ 1IT11m Ilf II t\I~111 \"1"[. ,11 t~ H11 ~i:;: MTh~ihhjH~H'ftlftti Ii.!, :1'
:!q !!.; il!! i ! ili; I 1111111 ill. !!\ "I' !1 LDl
0.2 " , ,I;; 1;;: i:l! !ii,: 11': :111 Illi Iii 1 !;i: \!!' il~.'lli ii 1111 :\\' From: Data Book on Hydrocarbons;
!:,: ! :;1:: i:!! :11; 11\1 11,11 i,illlll'I'I'!~III:I:i! ITI! ! Ii 1'1 ( II by B. J" Maxwell I
': ' I I: ,III ,;1, 'I!' II!\ 'III >I I I ,I 11! ' I 1
1m !i:: ifll! rl:f! 1 WI: rill" mr ill, Tjl /1 jnl:"! ili! 1'11'11 1.1' 111'1'1- ~~~::t~~st~:d ~~~:~.). Ir
'1'1 \'" I' 1 ' II 'II "II' '1'1 'I 'II 1,\1 I . ..
: o! I :! : '\ ' 'I : :: ! : I ! : 'II i, I ,I : :! I: 11:i I' !
:::; iil: liii Illi II Ii!! 1I11 :Iil !i!1 Iii! !Iillill III .1 I Ililllllllllliil!iilli:i!liiiil!iiili!::II!\ililiil !\IIIIIIIIIIII
0.1
o
-
c
Q)
u
>-
.-
'Vi
o
U
III
>
20
40
60
80
100
120 140 160
Temperature. F
180
200
220
240
260
280
FIGURE II.
COMPARISON OF LlaUID VISCOSITIES
-------
33
The aromatic components are somewhat higher in viscosity than the paraffins
but are less abundant in both gasoline and Stoddard Solvent.
The following
tabulation shows surface tension data for selected hydrocarbons and includes
the value for water as a reference. (4)
2 methylhexane
(iso-heptane)
2 methylheptane
(iso-octane)
Surface Tension
dynes/em at 72F
18.2
20.0
21.4
19.1
Hydrocarbon
n-hexane
n-heptane
n-octane
20.5
2,2,4 trimethylpentane
(iso-octane)
toluene
18.6
p-xylene
1,3,5 trimethylbenzene
28.1
29.5
28.2
27.8
28.5
29.3
72.6
o-xylene
m-xylene
1,2,4 trimethylbenzene
water
It is expected that the values for n-heptane and Stoddard Solvent
should be very close to that for gasoline.
A comparison between gasoline and Stoddard Solvent can be made on
the basis of distillation curves. Stoddard Solvent, like gasoline, is a grade
of petroleum distillate, but generally of lower volatility than gasoline. The
distillation curves for a typical gasoline (according to ASTM specification
D484-s2) and Stoddard Solvent (according to ASTM specification D-86) are com-
pared in Figure 12.
Stoddard Solvent can be considered as a high-boiling
fraction of gasoline and as such should have properties similar to those of
gasoline with the exception of vapor pressure.
The approximate composition of the Stoddard Solvent used in this study
was 51.4 percent (by volume) paraffins, 39.8 percent napthenes, 8.8 percent
aromatics, and 0 percent olefins. By comparison, a "typical" pump gasoline
-------
34
Stoddard solvent
400
300
~
Q)
~
:J
o 200
~
Q)
a.
E
Q)
t-
./1'
/.
./
100
o
IBP
20
40 60
Percent Evaporated
80
End
FIGURE 12. STANDARD DISTillATION CURVES FOR A
REPRESENTATIVE COMPOS ITE GASOLINE
AND FOR STODDARD SOLVENT
-------
35
would contain about 59 percent paraffins, 0 percent napthenes, 31 percent
aromatics, and 10 percent olefins.
From the comparisons between viscosity, density, composition, and
surface tension, we believe that both n-heptane and Stoddard Solvent are suit-
able substitutes for gasoline in studies of droplet impaction, liquid film
transport, and droplet reentrainment. We believe also that useful information
can be gained by measuring vaporization rates for these fluids during the drop-
let impaction studies.
Mixture Quality Measurement
"The requirements for the mixture quality measurements were primarily
to assess the ext~nt of impaction of fuel droplets on the walls of the test
sections and to provide information on the distribution of fuel droplets in
the cross section of the test sections. A secondary requirement was to allow
measurement of the distribution of fuel between the vapor and liquid phases.
Several sampling techniques were considered as possible alternatives
and two techniques were used.
The sampling technique used for the reported
measurements was a simple probe, which collected both drops and vapor at various
positions in the test section cross-sectional area, coupled with liquid skimmers
on the walls and liquid temperature measurements. Analyses of the collected
samples were attempted by use of a thermal conductivity cell in a gas chroma-
tograph, condensation with ~ubsequent weighing, a F.I.D. hydrocarbon analyzer,
and a combination of weighing and the hydrocarbon analyzer. The hydrocarbon
analyzer alone was used for the reported measurements.
Vapor/Droplet Samplers
Several vapor/droplet sampler designs were used in attempts to
separate liquid droplets from the mixture stream quickly in order to measure
fuel content as vapor and as liquid. The premise behind these designs was
that little vaporization of the liquid droplets could occur in a single-stage
impactor because the residence time would be short before the liquid was
separated from the air.
Further, if the liquid droplets could be impacted on
a surface from which they could be removed and rneasured, the remaining mixture
-------
36
would contain only vaporized fuel.
To accomplish this separation, several
different impactors having a fine screen or wire mesh as an impaction surface
were designed.
These impactors are designated as Designs 1, 2, and 3 and
are illustrated in Figures 13, 14, and 15.
A discussion of the procedures and principles used to design the
impactor samplers for this study is given in Appendix A.
The design shown in Figure 13 was tested with and without air flow
through the wire mesh impaction surface.
This design was abandoned because
there was found to be excessive loss of droplets to the walls of the probe
which, because of the apparatus geometry, were up to 6 inches in length with
one or two bends.
Furthermore, experimental checks and analytical estimations
indicated that if the fuel had not saturated the air stream at t4e point of
sampling there could be significant changes in the liquid-vapor distribution
of the fuel within the length of the sampling probes which in some cases were
almost one-half as long as the test section.
The second impactor design, shown in Figure 14, was abandoned because
of excessive loss of liquid to the wall on the vapor side within the impactor.
The cause of this loss was assumed to be that the impaction surface (screen)
was too small in diameter in comparison to the impactor jet diameter.
The third impactor design, shown in Figure 15, contained a larger
impaction surface which was intended to remedy the assumed cause of unsatis-
factory operation of the design shown in Figure 14. However, it was found that
the same problem still occurred with this design. The source of the problem
was traced to intermittent release of liquid deposited on the inner wall of
the probe tip which flowed along the wall and collected on the inside of the
impactor housing next to the jet itself., Release of drops from this location
caused liquid to be flung against the wall on the vapor side of the impactor.
The importance of this loss of liquid to the vapor side could not be quanti-
tatively assessed but it was decided that the measurements of vapor concentration
in the mixture could be seriously in error.
It should be noted that in short
duration testing and checkout of this design the loss of liquid to the vapor
side did not show up. The concentrations of liquid in the mixture passing
through the test section compounded the problem by causing liquid rates
into the impactor which exceeded the capability of the impactor to adequately
separate the liquid from the vapor. The flow rate through the liquid side
-------
37
g;
?\
I
I
~"
~..........
/
'j
j
-----..
-- - - - --. . - - -- - - - - -
\
---~_. - -_. .. -
Gn.- 5~,-9.1.~
\ -
I
\
--
\-- -- ---. ,
---1-
.~"
(~.
TE.:7T ...5c.<./70/'l--'
.-~~ \ 7,
I .
;'H ?<:;;!F-
c--.
I
.
L/Of //0
/7/VC-ff?E.::v7 ...5C,f""cE/v
I ..5i9.W'~#W$"~.: hCW'.$4'I-tS' - ~
t: ~--fl'O HE.K. Na C"HP .5C. X /- LC!i
.3 ...s;.vMPLI,M:7 ~8.: TV8E
-'I- C/i'LIB~HT70N .~)(
.5 ~~#'IG' /"7fV6e HOV-S/'/VG -L!fOT??:)M
~ "'-A"~"'~-O/O O:~AG' ,7/.0.,( o..a
T fflR'"t"E.e" "'~-O:.t:: O--~ ~,(o.x/ a
~ ...s;9'~r2:/.M:;' .~ ~A"Er .
----
9 CHL~HJ7O/V 40,,1' eND n~
10 ""O-~4 .#.D: -'To. CHP...5C. .x ~ L dO
II .5rO£L77NC:; £I::1V..!'TA'7L J'L/~ "::5;;--'400
..-.--------
/~ ...s;.v,wr'Z/AG ~£ A'Z'T.'9'/N'I.M:i' A"7M:i'
-~ . - --
E':: -. ,
.r: .,
.-c
:z
I
~
.....
".
0)-
FIGURE 13. VAPOR/DROPLET SAMPLING PROBE - DESIGN I
0V
'0)
-------
38
Vapor line
-- --11
I
Probe
I
I
---~
_...._~
---
Liquid line
,
.
- J
Sample probe
chamber
Traversing mechanism
FIGURE 14. VAPOR/DROPLET SAMPLlN<, PROBE - DESIGN 2
Screen
Liquid line
FIGURE 15. VAPOR IDROPLET SAMPLING PROBE - DESIGN 3
-------
39
had little effect on liquid loss to the walls of the impactor housing even
for rates through the liquid side which were as high as three times the rate
on the vapor side.
As was mentioned, the major problems with the vapor/droplet samplers
were liquid collecting along the outside 'of the jet plus high liquid loadings
in operation which in combination led to loss of liquid onto the vapor side.
attempt was made to circumvent these problems by modifying sampler Design 2.
A wire mesh screen was installed to replace the solid surface around the jet.
An
The liquid collecting on this screen was drawn off through a plenum behind the
screen and added to the liquid-side sample. This procedure worke~ fairly well
but was abandoned because of continued liquid leakages and losses into the vapor
side with this design resulting from the high liquid loadings.
It would be
instructive to attempt the same procedure with sampler Design 3, since for this
design the impaction surface is larger and more adequate for high liquid loadings
than that of Design 2, and removal of the liquid collecting around the jet holds
considerable promise of giving satisfactory operation. However, due to time con-
straints and continuing problems, all of the vapor/droplet sampler designs were
abandoned and an alternative technique using wall fuel-film skimmers and a total
droplet/vapor sampler was adopted.
The final sampling probe which was used in this study was a simple tube,
shown in Figure 16, which collected both the vapor and liquid simultaneously.
The probe tips were sized to give isokinetic sampling. When samples collected
in this manner were coupled with liquid temperature measurements and liquid
skimmer rates, calculation of the desired information on liquid distribution in
the test section could be made.
Wall Fuel~Film Skimmers
The major effort of the droplet-impaction.studies was to determine
the fraction of liquid as droplets that deposited on the walls of the test
sections from the mixture entering the test sections. The amount of liquid
deposited and remaining on the walls of the test sections was determined by
direct measurement by weight over measured time periods of the liquid removed
from the wall by fuel-film skimmers.
-------
40
The concept of a wall-skimmer design is based on the premise that
liquid flowing along the walls of the test sections can be removed by suction
through an opening in the wall. The wall section at the opening should be
designed to not obstruct the main mixture flow, and air and fuel vapor flow
through the opening should be minimized. To accomplish these objectives,
short skimmer sections were installed at the entrance and at the exit of the
test section.
The walls of these skimmer sections were constructed of porous
metal with an annular space behind them.
Suction on the annular- space caused
. air and liquid to be drawn through the porous metal. The operation of the
skimmers was controlled so that the air flow was no greater than that required
to completely remove the wall film.
In all cases the air flow through the
skimmer sections was in the range of 0.6 to 1.0 scfm.
4.5
~
o
T
0.250
70 scfm: 0 = 0.136
40 sc f m: 0 = O. 157
20 scfm: 0 = 0.209
I
I
-1r
0.125
F I GU RE
16.
SIMPLE TUBE SAMPLING PROBE
Figure 17 is a detailed drawing of the skimmer sections.
Flow from
the annular space behind the porous metal plates was from each of the four
sides of the square cross section. This allowed for liquid removal in any orien-
tat ion of the skimmer section.
The liquid and air removed from the skimmers
-------
41
passed through an all-glass impinger submerged in an ice-water bath. The
impinger also contained a glass fiber demister at the outlet so that droplets
entrained by the gas bubbling through the collected liquid would be retained
within the impinger.
A
t:;:\ 11111 t;\
I.:J '~ II I.:J
____~L__-,
o : 1-------' I 0
I 1 .1 I
I 1 I -I
==:',,-_J I I l__,.::=
I 1--
= =,- -I I I f - ,,= =
I I I I
I L_______J :
01
'- - --;tl-- _..J
~ 11'~
\:!:I II III
j 0.750"
-I typical 8
places
Drill 0.125'" diam
top for 1/8 II 28 machine threads
4 holes
A
1.437"
o typo 8 places
~-
No.9 (0.1960) drill
through 8 holes
Cross - Section
A-A
FIGURE 17. FUEL-FILM WALL-SKIMMER DESIGN
The wall-skimmer positioned at the entrance to the test section
assured that no previously deposited liquid film entered the test section.
second wall skimmer, positioned at the exit from the test section, collected
The
liquid deposited and remaining on the test section wall at this position. This
second skimmer then gave a direct measurement of the total amount of liquid
deposited and remaining on the wall of the test section. it can be shown that
a mass balance on fuel injected into the test section is possible with measure-
ments made on the injected fuel amount, the wall-skimmers, and the total vapor-
plus-droplet sample. This total mass balance however does not provide a measure
of the amount of fuel entering or exiting the test section as a vapor.
To
approximate the amount of vapor in the mixture, additional measurements in
the form of fuel temperature are required.
-------
42
Fu~~ 'y'~rization Measurements
The extent of fuel vaporization within the test section was measured
in a somewhat indirect method by determining the fuel temperature at the inlet
section and at the exit from the test section, and by calibration of the
equilibrium vaporization amounts for the fuel as a function of temperature.
The assumption inherent in using fuel temperatures as an indication of the
extent of vaporization is that the vaporized fuel is in equilibrium with the
liquid fuel. Experiments with the vapor/droplet sampler for short time periods,
before liquid leakage onto the vapor side of the sampler became ~roub1esome,
indicated that this assumption was correct.
Special experiments were also con-
ducted by introducing fuel into the system at different temperatures.
In these special experiments the ambient air temperature was 79 F,
the air flow rate was 70 scfm, and fuel was introduced onto the spinning disk
atomizer at 49 and 74 F.
The resulting fuel temperatures at the inlet to the
test chamber were nearly equal at 58.5 and 59 F, respectively.
Nearly adiabatic
conditions can be assumed for the passages between the spinning disk and the
test section, consequently, the nearly equal final fuel temperatures indicate
that equilibrium was reached at the inlet to the test chamber.
In the experimental,apparatus the liquid temperatures were measured
with thermocouples located in the wall fuel film at both the inlet and outlet
from the test section.
The thermocouple beads were positioned off the wall
surface, but within the liquid film. For a few tests a thermocouple was
attached alongside the sampling probe and measured the temperature of the air-
droplet flow. The results of these special experiments using different inlet
fuel temperatures and the short-time results with the vapor/droplet samplers
indicated that the vaporized fuel is in equilibrium with the liquid fuel at the
inlet to the test section. Knowledge of the amount of fuel vapor in equilibrium
with the liquid at different temperatures allowed fuel temperature to be used
as' an indicator of fuel vapor content in the air.
Calibrations were performed to determine the equilibrium content
of fuel vapor in air over liquid fuel. Because Stoddard Solvent. which was
the 'fuel used in the experiments, has some dis tribution of molecular weights
in its compositional makeup it is important that the amount of liquid used in
-------
j-
43
the experimental equilibrium determinations should be such that the fraction
evaporated during calibration is comparable to the fraction expected to evap-
orate in the test section. This effect was taken into account in two ways.
First, an estimate of the amount of fuel vaporized in the test section as com-
pared with the total amount of fuel injected was used as a guideline in choosing
the amount of fuel to be used in the equilibrium determinations. Secondly,
various amounts of fuels were used to note any effect that quantity might have.
The equilibrium determinations were performed by partially filling
a nylon bag with air and injecting the fuel into the bag in such a quantity
that there would be sufficient fuel remaining after equilibrium had been reached.
The bag was then allowed to remain at the temperature of interest for several
hours after which the amount of fuel in the air-fuel vapor mixture in the bag
was measured. The results of these equilibruum determinations are shown in
Figure 18.
These special procedures were used rather than the standard equilib-
rium-air-distillation (EAD) procedure because the primary objective of the
calibration was to determine the fraction of fuel evaporated under conditions
duplicating the actual test conditions as nearly as possible.
The ASTM distillation curve for Stoddard Solvent, shown on Figure
12, can be used to estimate the dew point temperatures for a range of air-
fuel ratios.
The dew point temperature is the temperature of the 100 percent
evaporated point of the EAD. This was done for air-Stoddard Solvent mixtures
using the Bridgeman(5) alignment charts. The resulting dew point temperatures
were: 125 F for a 16:1 air-fuel ratio, 104 F for a 30:1 air-fuel ratio. and
75 F for a 70:1 air-fuel ratio.
On most of the test runs. except for the
vaporization tests, the mixtur.e temperature was less than 80 F and generally
less than 75 F.
Consequently. it is possible in almost all cases to have
vaporized fuel in equilibrium with the liquid fuel.
Vapor Analysis System
The amount of fuel contined in a gas sample was determined by a
Beckman Hydrocarbon Analyzer employing a flame ionization detector.
Several
other techniques were attempted throughout the course of this study and included
a gas chromatograph with a thermal conductivity detector and a condensation
-------
44
20
2
,v
V
1/
Y 0
~
/
0 /
-" ~
- ~
-
18
rt)
E
0
.....
E
CI
:L
- 14
:2
:J
c:r
...J
s:: 12
::
~
-
0
0
-
c
0
U
c
.-
II)
0
C) 8
-
0
-
c
G)
- 6
c
o
u
G)
:J
~ 4
16
10
050
55
60
65 70 75 .
Temperature, F
80
85
90
FIGURE 18. EQUILIBRIUM VAPOR CONTENT IN AIR OVER STODDARD
SOLVENT AT ATMOSPHERIC PRESSURE
-------
45
technique which collected vapor from the gas sample for sub~equent weighing.
These latter two techniques were not found to be suitable for various reasons.
The gas chromatograph technique was intended to be a continuous
measurement of hydrocarbon content in the gas sample. In this case a steady
stream of the gas sample was passed through the thermal conductivity detector
in a gas chromatographic unit. It was found, however, that it was not possible
to control the pressure and temperature of the sampled gas stream with sufficient
precision so that the indicated hydrocarbon content could be accurately measured.
In addition, the heated tungsten wire of the thermal conductivity detector
oxidized badly in the gas stream and deteriorated rapidly with use. For these
reasons this technique was abandoned.
The second alternate technique employed a cold trap to condense fuel
vapor from the gas sample on a continuous basis for subsequent measurement.
There were two main problems with this technique. First, water vapor in the
ambient air also condensed in the cold traps making mass measurements of the
fuel difficult. It would have been necessary. to design and build a dryer for
all the incoming air to the test section to conveniently avoid this problem and
it was decided that such an approach was impractical. A second problem with
the condeasation scheme was the time required to obtain a measurable sample. A
mass analysis by weighing the collected liquid was felt to be the most direct
technique; however, the sampling times to give a weighable sample were excessive.
The technique employed in the study using the hydrocarbon analyzer
was direct and convenient. A gas sample was drawn from the test section through
the sampling probe into a nylon holding bag from which the sample was then passed
into the hydrocarbon analyzer. This sampling bag technique was necessary
because the hydrocarbon analyzer can only be operated at ambient pressures,
whereas, sampling had to be done at the test section vacuums. Adjustment to
ambient pressures was possible after a sample had been collected in ~he bag.
Gas Sampling System
The gas sampling system was designed to collect both vapor and
droplets in the mixture stream at the exit from the test section. The sampling
system is shown in Figure 19 and consisted of a sampling probe positioned at
-------
46
the exit from the test section with the,capabi1ity of traversing or being moved
in the test section cross-sectional aLea. The sampling flow rate was fixed at
.,
20 liters/min (0.071 scfm) through the sampling nozzles which were sized to
sample at the average air velocity through the test section, even though the
flow was pulsing and sampling could not be isokinetic.
Vacuum pumps were used to
Test
section
~
~
-. -- Sampli ng
probe
, '.
Reference
gas
R
Critical flow - -
orifice
+
Vent
Vacuu m
pump
FIGURE 19.
FLOW DIAGRAM OF SAMPLING SYSTEM
draw the sampling flow through a primary sampling line where the flow rate
was monitored with. a rotameter and with temperature and pressure measurements
for flow correction. The sample for analysis w~s taken from this primary
line into the sample bag at a rate of 700 cm3/min (0.025 scfm) as controlled
by a calibrated critical flow orifice. '
-------
47
The nylon sample bag was contained within a metal canister inside
an oven maintained at 150 F. Sample was drawn into the bag by pulling a
vacuum on the metal canister in which it was contained.
Samples were
collected for analysis over time periods of approximately 5 minutes and at
all other times the flow through this take-off line was bypassed around the
sample bags so that this take-off flow rate was maintained at 700 cm3fmin.
All sample lines leading from the sample probe through the rotameter. the
samplp. bag. and the critical flow orifice. and from the sample bag to the
hydrocarbon analyzer were maintained at about 200 F. The analysis of the sam-
ple bag was performed by drawing a sample from the bag through the hydrocarbon
analyzer.
The rotameter on the primary sample line was calibrated to provide
a correction factor as a function of temperature and pressure of the gas
entering the rotameter.
These correction factors in conjunction with the
calibration curve for the rotameter at standard temperature and pressure
permitted calculation of the flow through the rotameter based on standard
conditions. The flow into the sampling nozzle was then obtained by correcting
the staridard flow for the conditions which were present in the test section.
The standard calibration of the rotameter and the experimentally determined
correction factors are presented in Appendix B. Figures B-15 and B-16.
The analysis of the hydrocarbon content of the gas sample collected
in the nylon bag was obtained with the hydrocarbon analyzer as calibrated
with Stoddard Solvent. The calibration of the hydrocarbon analyzer is pre-
sented in Appendix B, Figure B-17. This calibration was performed by injecting
predetermined amounts of Stoddard Solvent into a nylon bag which was then filled
3
The volume of the calibration bag was 3960 cm
this volume by measuring the gas pressure in the
with air to a known volume.
(0.14 ft3) and was filled to
bag and filling it until the gas pressure was the same in all cases. The in-
jected fuel was allowed to vaporize while the sample bag was contained in the
oven maintained at 150 F. After the injected fuel was completely vaporized.
a sample was drawn from the bag through the hydrocarbon analyzer and the
reading noted.
It was found that the gas pressure at the hydrocarbon analyzer had
some effect on the indicated readings. For this reason the pressure of the gas
in the hydrocarbon analyzer was monitored with a water manometer.
Calibration
-------
48
on a daily basis was performed with the pressure being adjusted for each daily
calibration such that the desired reading for a known hydrocarbon source was
obtained.
This pressure was then maintained for all measurements taken on
that day. It was also found that the temperature of the hydrocarbon analyzer
had a small effect on the indicated readings but more important, if the temper-
ature of the hydrocarbon analyzer was sufficiently low, such as room temperature,
adsorption of the fuel vapors could occur within the analyzer.
To avoid this
complication the hydrocarbon analyzer was maintained at a temperature of about
135 F through the tests.
This temperature was found sufficient to maintain
the hydrocarbon in the gas stream without appreciable adsorption or condensation
within the instrument. This was experimentally determined by maintaining the
hydrocarbon analyzer at various temperatures until a level was found which did
not lead to hold-up by adsorption within the hydrocarbon analyzer as determined
by subsequent desorption indicated by the detector in the hydrocarbon analyzer.
Owing to the possibility that adsorption of hydrocarbons from the
sample gas stream onto the inner surfaces of the sampling lines could occur
to varying degrees depending on temperature of the sampling line, experiments
were performed to determine the importance of this factor. The experiments
consisted of operating the sampling system with the sampling lines at room
temperature and again at 150 F, and for each temperature drawing an air-vapor
sample of the same concentration into the hydrocarbon analyzer. It was found
that for air saturated with Stoddard Solvent at room temperature there was no
appreciable loss by adsorption of the hydrocarbon on the internal surfaces of
the sampling lines even if the lines were maintained at room temperature. How-
ever, to avoid such problems the sampling lines were maintained at elevated
temperatures.
Experimental Program
The experimental program was designed to provide information on
droplet impaction in simulated intake manifold sections as a function of
drop size, air flow rate, manifold geometry, mixture temperature, mixture
pressure, and surface roughness.
Further, the characteristics of liquid-film
flow along the walls of the test sections were to be determined in terms of
liquid-flow patterns and the amount of liquid carried along the wall.
Fuel
-------
49
vaporization was also investigated as a function of inlet-air temperature
and temperature of the test section walls.
Laboratory System Limitations
The laboratory system was limited in several respects. First, the
flow rate which could be achieved through the test sections was dependent on
whether steady or pulsed flow existed and the pressure level desired in the
test section. The capabilities of the air flow in this regard are illustrated
in Figure 4.
Another serious limitation of the experimental apparatus was
concerned with the vaporization of the liquid fuel as it flowed to the atomizer.
Vacuums of greater than 15 inches of mercury resulted in serious formation of
bubbles within the fuel lines. This was controlled somewhat by the fuel being
chilled in an ice-water bath prior io its passage through the control valve where
bubbling most often occurred. This cooling in conjunction with system operation
at vacumms less than 12 inches of mercury ensured operation without vaporization.
Experimental Conditions
The experimental conditions were held constant to the extent possible
with the exception of the variables being studied. All experiments were run
with the inlet air at ambient temperature, the liquid fuel at about 50 F, an
air flow pulse rate of 1000 cycles per second, and a fuel injection rate of
0.344 1b/min. Temperatures of the sampling and sample transfer lines were
. .
maintained at about 200 F and of the oven at 150 F. The sampling and sample.
transfer lines were heated with electrical heating tape and temperatures at
points along the lines were measured with thermocouples. The hydrocarbon
analyzer was kept at 135 F.
The variables consisted of air-flow rate, fuel-droplet size, and
test-section geometry.
Mixture samples were taken from at least three loca-
tions in each test section cross-sectional area.
Figure 20 shows the sample
probe locations for the different test runs. A sample taken from one probe
position constitutes a test run. A matrix of the variable test conditions is
given in Table 2. From the matrix it is seen that almost 80 test runs were
recorded.
Over 200 test runs were made prior to these recorded test runs during
-------
so
0.50 (typ)
0.25 (typ)
O ",
( )
4 '-"9
I
01
I
""6
()
I
~5 () 0
V -7 2
I
""
I \
""'/8
J
03
.. ~06~5 -j I
'~1.25~
FIGURE 20. CROSS-SECTION OF MANIFOLD TEST
SECTION SHOWING PROBE POSITIONS.
VIEWED FROM OUTLET
-------
51
TABLE 2.
MATRIX OF TEST CONDITIONS
System Air Flow
Test Section 70 scfm
Geometry 20 scfm 40 sefm (1) (2)
Entrance to
Test Sections 2,5,4* 2,5,4 2,5,4 2,5,4
Straight 2,5,4 2,5,4 2,5,4 2,5,4
30° 1,5,3 1,5,3 1,5,3 1,5,3
90° 4" R 1,5,3 1,5,3 1,5,3 1,5,3
90° 0" R 1,5,3 1,5,3 1,5,3 1,5,3
90° 4" R(a) 1,2,3
4,5,6
7, ,9
90° 4" R(b) 1,5,3
900 4" R(c) 1,5,3
900 4" R(d) 1,5,3
Disk speeds: 20 sefm: 55,000 rpm 70 sefm (1): 55,000 rpm
40 sefm: 55,000 rpm 70 sefm (2): 26,000 rpm
11" Hg
8.5" Hg
2" Hg
Vacuums in test section:
(All tests at pulsed flow)
20 sefm:
40 sefm:
70 sefm:
* Probe positions -- See Figure 20
(a) Replication with more detailed traverse
(b) Rough-wall test section
(c) Aluminum test section -- heated inlet air
(d) Aluminum test section -- heated walls
-------
52
checkout and preliminary studies. Most of these were concerned with solving
problems occurring with the vapor/droplet samplers, mixture generator, hydrocarbon
analysis technique and sampling flow system.
lost because of leaks in the sampling bags.
In addition, a number of runs were
The experimental conditions were chosen to simulate, to the extent
practical, the flow, pressure, and velocity in an intake manifold.
The pressure
level in the test sections at the lowest flow rate is somewhat higher than
would be encountered in an actual induction system at a comparable air flow rate
because of the vaporization-bubbling problem previously mentioned.
The experi-
mental apparatus was operated at 11 in. Hg vacuum for 20 scfm air flow while
the vacuum expected in an induction system for this flow would be about 16 in.Hg.
Another deviation from an exact simulation for an induction system is
in the air/fuel ratio entering the test section.
The fuel atomized in the mix-
ture generator at the set rate of 0.344 lb/min corresponds to air/fuel ratios
of 18:1 at 70 scfm air flow down to 5:1 at 20 scfm. However, because of the
considerable loss to the wall in the mixture generator, the air/fuel ratios
entering the test sections were of the order of 30 to 70.
This is probably not
a significant factor in a study of deposition rates but would be significant in
studies of reentrainment after deposition.
Drop Sizes Entering the Test Section
Another test condition of importance is the size distribution of the
drops entering the test sections.
This was measured with an impaction-slide
scheme. A drawing of the device specially constructed for this purpose is
shown as Figure 21. This drop sizing device is designed to fit into the labor-
atory flow system just upstream of the test sections. Since the impaction sur-
face disrupted the air and fuel-droplet flow, the sizing apparatus was installed
only for the purposes of collecting drops and at all other times was removed.
The principle of operation is that the air flow in the system cross-
sectional area is sufficient to give deposition into a slot on the traversing
slide and allow the droplets to impact onto the coated glass slide. Calcula-
tions based on impaction into recessed bodies(6) maintained in a free air
stream (velocities taken as averages for test conditions) indicate that the
-------
-._--
Mixture
flow
'\7
I
I
I
I
I
I
I
I
I
I I
I I
- .1": --=- -= =- -=-=- -=-:j = -
,-I...r-- ---- -- -=- -:... -=. -:LI:J
-1- - .- .- - - - - -I
-r------,-
'''~'\'-A ,,,,\,\,\,\
I ! : I
~""\f~"- - - - - '~'fl'."
I
I
1
I
I
I
I
I
I
1
53
~ Traversing slide
,,--Coated gloss slide
~ Gloss-slide support
~Base for
traversing slide
FIGURE 21.
COLLECTION APPARATUS FOR DROP SIZING
-------
54
impaction device should be adequate to collect droplets about 5 microns and
larger.
In practice the system was more effective than predicted.
The glass
slide impaction surfaces were coated with magnesium oxide particles into which
the droplet impacted, leaving a crater.
The fuel-air flow was first allowed to stabilize while the traversing
slide was kept in a position such that the opening was outside the apparatus.
Sampling was then accomplished by passing the traversing slide with the sampling
opening across the flowing gas stream. As the opening or slot passed across,
drops were impacted by their inertia through the opening onto the coated slide.
The slide coating was a loose agglomerated mass of magnesium oxide
smoke particles. This coating was penetrated by each drop as it impacted on
the slide, leaving a crater in the coating as an indication of drop size.
Extensive calibration of this technique has indicated that the true drop sizes
are 0.81 times the crater size(5). Size distributions were determined with an
optical microscope.
The size distributions of the drops entering the test sections at
different gas flow rates are shown in Figure 22.
All size distributions
given are averages across the test-section cross-sectional area as provided by
the slide movement.
The disk speed for air flow rates of 20, 40 and 70 (1) scfm ..
was 55,000 rpm. For the case noted 70 (2),the speed was 26,000 rpm. As is seen
there is little difference among the size distributions. On a mass or volume
basis there would be some differences because the case 70 (2) indicated a number
of very large (~250~) drops which would significantly shift a mass mean to
larger sizes.
General Test Procedures
The procedures employed in performing the tests were quite direct
and, except for setting up the primary sampling rate, involved only adjustment
to prescribed values. The primary sampling rate was fixed by an iterative
process involving calculated temperature and pressure corrections with sub-
sequent flow adjustments until the sampling rate was 20 liters/min (0.706 sefm).
The general procedure was to first allow all controlled temperatures
to be brought to the desired levels. The temperatures of interest in this
regard are the sampling lines (200 F), the oven (150 F), and the hydrocarbon
-------
55
99
98
80
1 '
Volumetric Flow Rate for Gas
/ V 20
/ 40 and 70 (I)
/
y/ ~ .... 70 (2)
~"
I I>'"
-:.'
If .'V
i/~'"
f
/,1
~!:I
,'. '/
.?/!/
JI
l V '
~
/ II
J Airf low, Disk Speed,
/: h( scfm rDm
400'
I) --0-- 20 55,000
..,j f70( I) --~-- 40 55,000
I
/f -'-6-' - 70 (.U 55,000
201 / :'70(2)
.. ., ''\]..'" 70 (2) 26,000
95
90
L- 70
a>
~ 60
::J
Z 50
>-
..0 40
-
c
~ 30
L-
a>
a. 20
a>
>
-
C
::J 10
E
::J
U 5
2
1
0.5
0.2
0.1
0.05
0.01
2
3
4
6 8 10 20 30 40
Equivalent Particle Diameter, microns
60
80 100
FIGURE 22. SIZE DISTRrSUTIONS FOR DROPS ENTERING
TEST SECTIONS
-------
56
analyzer (135 F).
After the proper temperatures had been achieved, the main
air flow, pu1ser motor, and skimmer flows were started.
Pulse rate was adjusted
to 1000 cps and then air flow rate and pressure in the test section were set at
the prescribed values. At this point the primary and take-off sample flows
were started and adjusted.
After proper adjustment of all conditions and flows had been achieved,
the spinning disk rotation rate was set at either 26,000 or 55,000 rpm. At
this point the fuel flow was started and timing for Skimmers No.1 and No.2
was begun. Fuel flow through the mixture-generator chamber skinuner was deter-
mined once or twice over timed intervals during the course of three test runs.
Weight measurements over time intervals gave chamber skimmer rates.
Weight
measurements were also used to give rates at Skimmers No.1 and No.2, but the
time interval was that for the three test runs.
Three test runs were run sequentially each time the flow system was
operated. Each test run consisted of a gas sample taken at one location in
the test section exit. Each test or gas sample was initiated by directing
the take-off sample flow into the sampling bag after stopping the bag by-pass
"
flow. Sample was drawn into the sample bags for about 5 minutes at a rate of
700 cm3/min (0.025 scfm) until pressure measurements indicated the bag was full.
The by-pass flow was again started, the can containing the sample bag vented
to the atmosphere and the sample directed through the hydrocarbon analyzer.
After the sample had been analyzed the sample bag was completely collapsed and
the lines to the bag evacuated to about 20 inches Hg vacuum in preparation for
sampling from a new probe position.
Experimental Results
Experimental data obtained are presented in Tables 3 through 16.
These data were all generated using the simple-tube probe, and earlier data
obtained using probes with vapor/~plet impaction separators are not included.
The data presented can be summarized as follows:
Tables 3 and 4
Tables 5 and 6
Tables 7 and 8
- Test-section inlet
- Straight test-section
- 30-degree-bend test-section
-------
57
Tables 9 and 10 - 90-degree bend test-section with 4-inch bend radius
Tables 11 and 12 - 90-degree-bend test section with O-inch
bend radius
Tables 13 and 14 - Rerun of one test condition with more
detailed traverse
Tables 15 and 16 - Rough-wall and heated test section.
The significance of the values presented in the above tables is generally self-
evident, with some exceptions which require explanation.
The values for sample fuel concentration given in Column 13 of the
odd-numbered tables above are for the total fuel concentration, liquid plus
vapor, observed in the fuel-air samples collected by the traversing probe.
The values for local total fuel rate, given in Column 22 of the even-numbered
tables, are obtained by multiplying the observed total fuel concentration
(Column 13) by the volumetric air-flow rate (Column 3) and the appropriate
conversion factors. Thus, the local fuel rate to the probe is given in terms
of the total fuel rate for the test section that would be consistent with the
local fuel concentration.
The average fuel rate for the test section (Column 25)
can therefore be taken as the average of the local fuel rates (weighting the
local fuel rates according to probe position).
The fuel vapor rate to the probe (Column 23) is computed by multi-
plying the vapor concentration at saturation corresponding to the observed
test-section-exit wall fuel-film temperature by the air-flow rate.
amount of fuel in vapor form is computed and not measured directly.
Thus, the
This
procedure was necessitated by the difficulties in designing an effective
vapor/droplet separating probe, as discussed previously. The amount of fuel
present as entrained droplets at any probe position is computed as the difference
between the observed total fuel rate sampled and computed fuel vapor rate.
Experimental Error Analysis
Mass Balance.
,From the experimental data, the fuel rate both entering
and leaving the test section can be computed to provide a mass-balance check
on the accuracy of the data.
The flow rate to the mixture generator was the
-------
58
TABLE 3. EXPERIMENIAL DAIA FOR TEST-SECTION ENTRANCE
-- .-- -- .._---
,--_. _. ..-- -----.
2 3 4 6 7 8 9 10 11 12 13
Mesn Orifice Test Section Fuel Rste Chamber No. 1 No. 2 Samp le
Drop Air Chamber Air Wa 11 Film to Mixture Skimmer Skimmer Skimmer Fuel
Run Size. Flow, Vacuum, Probe Temperature, Temperature. F Generator, Flow, Flow, Flow, conc.;,1
No. microns scfm in. IIg Position F Inlet Outlet Ib/hr Ib/hr Ib/hr Ib/hr i>g/c
250 14 20 11.0 Rear (2)** 77 66.5 20.6 17.0 16.1
251 Center (5) 76 66.0 0.19 15.5
252 Front (4) 74 65.0 17.9 14.8
226 14 40 8.5 Rear 76 78.5 20.6 15.8 16.4
227 Center 79 " 0.25 17.7
228 Front 79 77.5 16.7 16.7
276 14 70 2.0 Rear 66 57.0 20.6 11.5
275 Center 66 57.0 " 15.7 0.14 14.8
274 Front 66 5 7.5 9.5
271 14* 70 2.0 Rear 65 58.0 20.6 11.4
272 Center 66 57.0 15.3 0.12 15.2
I 65 57.0 8.5
273 Front
.:..,-=-"~~:=_==--=::=--=-.:....:-= -..-
* Pills large drops. ** SC~ Figure 20.
TABLE 4. HEDOCEIJ DAIA FOR TEST-SECTION ENTRANCE
II, 1 '> 110 17 18 19 20 21 22 23 24 25 26 27
Test Loca 1 Local Average
Section Test Tota 1 Entrained Total Wall Total
Ai r Net Fuel Air Sactiqn Fuel Vapor Liquid Fuel FLlm SlIIIIpled
Mass toTes t Ent rained Saturated Velocity, Air Rate to Fuel Fuel Rate to Fuel Fuel
Run Flow. Section. Air/Fuel Air/Fuel average DenSit~, Reynolds Probe, Rate, Rate, Probe, Rete, RAte,
No. lb/hr Ib/hr Ratio Ratio ft/sec lb/ft. No. lb/hr lb/hr lb/hr lb/hr lb/hr lb/hr
250 3.4 26 115 49 .0469 1.20 0.78 0.42
251 90 118 1. 17 0.77 0.40
252 2.5 36 121 1.11 0.74 0.37
226 4.5 40 79 87 .0530 2.48 2.29 0.19
227 180 2.67 2.26 0.41
228 3.6 50 81 2.51 2.22 0.29
276 150 113 .0715 2.77 2.11 0.66
275 315 4.8 66 150 3.92 2.11 1.81
274 150 2.50 2.11 0.39
271 148 113 .0715 3.00 2.13 0.87
272 315 5.2 61 150 4.00 2.11 1.89
273 150 2.25 2.11 0.14
-------
59
TAJH.E 5. EXPERr:1EHTAL DATA FOR STRAIGHT TEST SECTION
2 4 6 7 8 9 10 11 12 13
H~an Ori Hce Test Section Fuel Rate Chamber No. 1 No.2 Sample
Drop Air (''hamb~ r Air Wall Film To Mixture Sk imme r Skimmer Sk immer Fuel
Hun Sixt:, Flow, Vacuum, Probe.: Tempera turt:) Temperature. F Generator, Flow. Flow, Flow, Cone.
110. microns scfm in. IIg Position F Inlet Outlet 1b/hr 1b/hr 1b/hr 1b/hr u.g/cm'3
24'J 14 20 11.0 R~ar 80 67.5 69.5 20.6 17.3 12.6
248 Ce:nte:f 80 0.16 0.15 13.0
247 front 80 68.5 68.5 17.5 13.0
225 14 40 8.5 Hear 71 70.5 71.5 20.6 16.5 15.2
224 Ce:ntef 71 70.5 70.5 0.27 0.27 15.2
223 Front 71 70.5 70.5 17.1 15.2
265 14 70 2.0 Rear 66 60.0 61.5 20.6 14.8 13.3
266 Ct:nte:f 65 58.5 60.5 0.07 0.37 13.3
267 Front 58.5 60.5 16.0 12.2
270 14", 70 2.0 Rear 66 57.5 59.5 20.6 13.6
26Y Center 66 57.5 59.5 14.8 0.07 0.35 12.6
268 Front 66 57.5 59.5 10.7
" Plus large drops.
TABLE 6. REDUCED DATA FOR STRAIGHT TEST SECTION
14 15 16 17 18 19 20 21 22 23 24 25 26 27
Test Local Local Average
Section Test Total Entrained Total Wall Tots
Air Net Fuel Air Section Fuel Vapor Liquid Fuel Film Sampl
Mass to Tes t Entrained Saturated Velocity, Air Rate to Fuel Fuel Rata to Fuel Fuel
Run Flow, Section, Air/Fuel Air/Fuel average Densit~, Reynolds Probe, Rate, Rate, Probe, Rate, Rat.
No. 1b/hr 1b/hr Ratio Ratio ft/sec 1b/ft No. lb/hr 1b/hr lb/hr lb/hr lb/hr lb/h
249 0.95 0.86 0.09
248 90 3.1 29 111 48 .0475 19,400 0.98 0.85 0.13 0.97 0.15 1.12
247 0.98 0.82 0.16
225 2.29 1.84 0.45
224 180 3.6 50 102 86 .0536 39,300 2.29 1. 77 0.52 2.29 0.27 2. 5f
223 2.29 1.77 0.52
265 3.49 2.34 1.15
266 315 5.2 61 146 113 .0715 68,800 3.49 2.29 1.20 3.42 0.37 3.7!
267 3.22 2.29 0.93
270 3.57 2.21 1.36
269 315 5.8 55 150 113 .0715 68,800 3.32 2.21 1.11 3.25 0.35 3.6!
268 2.82 2.21 0.61
-------
60
TABLE 7. EXPERIMENTAL DATA F~R 30-DECREE-BEND SECTION
2 4 6 7 8 9 10 11 12 13
Mean Orifice Test Section Fuel Rate Chamber No. 1 No. 2 Sample
Drop Air Chambe r Air Wa11 Film To Mixture Skimmer Skionner Skinuner Fuel
Kun Size, Flow, Vacuum J Probe Temperature, Temperature. F Genera tor, Flow, Flow, Flow, Conc.
No. mi.crons scfm in. IIg Position F In Ie t Outlet Ib /hr Ib/hr Ib/hr Ib/hr "g/em3
'2':'6 14 20 11.0 Top (1)*** 81 67.0 20.6 17.2 13.0
245 " " Center (5) 80 57.0 69.0 " 0.12 0.24 13.3
244 Bo ttom (3) 68.0 18.0 13.6
216 14 40 8.5 Top 75 68** 70*" 20.6 14.8
215 " Center 75 68** 70** " 15.8 0.20 0.39 16.1
214 Bottom 75 68~'d; 70}'r;'c 18.3
2~0 14 70 2.0 Top 68 63.5 65.5 20.6 9.3
281 " Center 68 63.5 65.5 " 15.6 0.08 0.71 10.1
182 Bottom 69 63.5 65.5 14.8
279 14'-' 70 2.0 Top 67 62.5 65.5 20.6 8.2
278 " Center 66 62.5 64.5 " 15.6 0.06 0.77 9.2
277 Bottom 66 62.5 64.5 14.2
* Plus large drops. *!Ir* See Figure 20.
** Approximate value.
TABLE 8. REDUCED DATA FOR 30-DECREE-BEND TEST SECTlON
14 15 16 17 18 19 20' 21 22 23 24 25 26 27
Test Local Local Average
Section Test Total Entrained Total Wall Total
Air Net Fuel Air Section Fuel Vapor Liquid Fuel Film Sampled
Mass to Test Entrained Saturated Veloci.ty, Air Rate to Fuel Fuel Rate to Fuel Fuel
Run Flow, Section, Air/Fuel Mr/Fuel average Dens it3' Reynolds Probe, Rate, Rate, Probe, Rate, Rate,
No. Ib/hr Ib/hr Ratio Ratio ft/sec Ib/ ft No. Ib/hr Ib/hr 1b/hr 1b/hr 1b/hr 1b/hr
246 0.98 0.85 0.13
245 90 3.0 30 114 48 .0475 ]9,500 0.99 0.85 0.14 0.99 0.24 1.23
244 1.02 0.85 0.17
216 2.22 1. 75 0.47
lI5 l/W 4.b 39 110 86 .0538 39,400 2.41 1. 75 0.66 2.45 0.39 2.84
214 2.75 1. 75 1.00
2~0 2.43 2.63
281 315 5.0 63 120 114 .0707 68,700 2.65 2.63 0.02 2.90 0.71 3.61
2~2 3.88 2.63 1.25
'279 2.15 2.63
278 315 5.0 63 124 114 .0707 68,700 2.41 2.55 2.67 0.77 3.44
277 3.72 2.55 1. 17
-------
61
TABLE Y. EXPEHHIENTAL DATA FOR 90 DEGREE-BEND, 4-INCH RADIUS TEST SECTION
3 4 6 7 8 9 10 11 12 13
:!cdn Oei flce Test Section Fuel Rate Ch ambe r No. 1 No. 2 Sample
IIrop Air Chumber Air Wall Film To Mixture Skimmer Skimmer Skimmer Fuel
Hun Sb:c, Flo\..., Vacuum, Probe Temperature J Terr.perature. F Generator. Fla.', Flow, Flow, Conc.
Nu. micruns scfm 1.11. IIg Position F Inlet Out Ie t lb/hr lb/hr 1b/hr 1 b /hr "g/cm3
24 J 14 20 11.0 Top 76 68.5 70.5 20.6 17.3 13.3
2!j ~ Center 0.18 0.21 13.6
2!.J Bottom 77 68.5 71.5 17.6 15.2
231 l!. !.o 8.5 Top 79 75.5 74.5 20.6 17.7 14.9
2JU Center 79 76.5 74.5 0.15 0.76 14.9
:!2Y Bottom 79 78.0 73.0 16.8 15.5
286 l!, 70 2.0 Top 70 64.5 65.5 20.6 9.8
~87 Center 70 64.5 66.0 15.7 0.10 1.14 9.5
288 Bottom 70 10.0
285 14", 70 2.0 Top 70 63.5 64.5 20.6 8.2
284 Center 70 63.5 65 " 15.9 0.06 1.01 8.1
283 Bottom 69 64.0 65.5 8.5
,', Plus larg~ drops.
TAIILE 10. REDUCED DATA FOR 90-DEGREE-BEND, 4-TNCII-RADIUS TEST SECTION
14 h 1(, 17 18 19 20 21 22 23 24 25 26 27
Test Local Local Average
Section Teat Tota 1 Entrained Total Wall Total
Air Net Fue 1 Air Section Fuel Va par Liquid Fuel Film Sa1\lpled
Mass to Test Entrained Saturated Velocity, Air Rate to Fuel Fuel Rate to Fuel Fuel
Run Flow, Section, Ai r/Fue 1 Air/Fuel average DenBit~, Reynolds Probe, Rate, Rate, Probe, Rate. Rate,
No. lb/hr 1b/hr Ratio Ratio ft/sec lb/ft 'Nm. lb/hr lb/hr lb/hr lb/hr lb/hr lb/hr
241 0.98 0.89 0.09
242 90 3.0 30 109 49 .0474 19,800 1.02 0.90 0.12 1.04 0.21 1.25
243 1.14 0.91 0.23
231 2.23 2.01 0.22
230 180 3.2 50 84 87 .0532 39,400 2.23 2.01 0.22 2.25 0.76 3.01
229 2.32 1. 93 0.39
286 2.57 2.63
287 315 4.8 66 124 114 .0705 68,400 2.50 2.69 2.55 1.14 3.69
288 2.62 2.69
285 2.15 2.55
284 31) 4.7 67 128 114 .0705 68,400 2.11 2.61 2.15 1.01 3.16
283 2.23 2.63
-------
6~
TABLE ll. EXPERUIENTAL DATA FOR 90-DECREE BEND, O-INCH-RADIUS TEST SECTION
2 4 6 7 8 9 10 11 12 13
Nean Orifice Test Section Fuel Rate Chamber No. 1 No.2 Sample
Drop Air Chamber Air Wall Film To Nixture Skimmer Skimmer Sk immer Fuel
Run Size, FImy', Vacut.m1, Prohe Tempera ture, Temperature I F Generator, Flow, Flow, Flow, Conc.
No. microns scEm in. IIg Position F In Ie t Outlet Ib/hr lb/hr Ib/hr Ib/hr ..g/cm3
240 14 20 11.0 Top 70 70.5 69.0 20.6 16.8 11.4
~J9 Center 70 70.5 69.0 0.20 0.26 10.7
238 Bottom 70 17.8 9.8
232 14 40 8.5 Top 78 78.5 75.0 20.6 15.6 14.9
233 Center 79 " 0.28 0.45 15.2
234 Bottom 79 77.5 75.0 16.2 18.3
292 14 70 2.0 Top 69 63.5 65.5 20.6 9.8
293 Center 70 63.5 65.5 14.8 0.05 0.95 9.8
294 Bottom 12.3
291 14" 70 2.0 Top 70 20.6 9.5
290 Center 70 63.5 65.0 15.7 0.05 1.01 9.5
289 Bottom 70 65.5 65.5 12.3
* Plus large drops.
TABLE 12. REDUCED DATA FOR 90-DECREE-BEND, O-INCH-RADIUS TEST SECTION
14 15 16 17 18 19 20 21 22 23 24 25 26 27
Teat Local Local Average
Section Test Total Entrained Total Wall Total
Air Net Fuel Air Section Fuel Vapor Liquid Fuel Film Sampled
MatiS to Tesl Entrained Saturated Velocity, Alr Rate to Fuel Fuel Rate to Fuel Fuel
Run Flow, Section, Air/Fuel Air/Fuel average Densit3" Reynolds Probe, Rate, Rate, Probe, Rate, Rate,
No. Ib/hr Ib/hr Ratio Ratio ft/sec Ib/ft No. Ib/hr Ib/hr Ib/hr Ib/hr Ib/hr Ib/hr
240 0.85 0.85
239 90 3.0 30 102 49 .0474 19,800 0.80 0.85 0.80 0.26 1.06
238 0.73 0.85
232 2.23 2.06 0.17
233 180 4.4 41 80 87 .0530 39,300 2.28 2.06 0.22 2.38 0.45 2.83
234 2.74 2.06 0.68
292 2.57 2.64
293 315 5.8 54 127 114 .0705 68,500 2.57 2.64 2.73 0.95 3.68
294 3.22 2.64 0.58
291 2.49 2.61
290 315 4.9 64 127 114 .0705 68,500 2.49 2.61 2.67 1. 01 3.68
2B9 3.22 2.64 0.58
-------
63
TABLE 1).
EXPERl~ffiNTAL DATA FOR REPLICATION TEST WITH 90-DEGREE-BAND,
4 - INCII-RADIUS TEST SECTION
I. 5 6 7 8 9 10 Il 12 13
Nenn Orifice Test Section Fuel Rate Chamber No. I No.2 Sample
Drop Air Chamber Air Wa Il Film To Mixture Skimmer Skimmer Skimmer Fuel
Hun Size, Fl 01', Vacuum, Probe Tempera ture I Tempera ture. F Generator. Fl 01< . Fl 01< , Fl 01<, Cone.
No. microns scfID in. IIg Position F In Ie t Outlet Ib/hr Ib/hr Ib Ihr lb/hr ..8 I cm3
298 14 40 8.5 Top (1). 73 75 20.6 15.7 0.03 0.81 16.1
:!Ij'} Top center (6) 73 75 16.4
296 Center (5) 73 75 16.1
~97 Bottom (8) 73 75 15.5
center
2Y5 Bottom (3) 73 75 17.5
JO! East (4) 74 76 17.7
303 East cented9) 77 79 16.4
302 Wesl cented7) 74 76 15.8
300 \~est (2) 74 76 16.4
. Se~ Fil;lIrc 20.
TABLE 14. REDUCED DATA FOR REPLICATION TEST WITH 90-DEGREE-BEND. 4-rNCH RADIUS TEST SECTION
14 15 16 17 18 19 20 21 22 23 24 25 26 27
Te8t Local Local Average
See t ion Test Total Entrained Tots I Wall Total
Air Net Fuel Air Section Fuel Vapor Liquid Fuel Film Sampled
Mass toTes t Ent rsined Saturated Velocity, Air Ra te to Fuel Fuel Rate to Fuel Fuel
Run Flow. Section, Air/Fuel Air/Fuel average Oe081t3' Reynolds Probe, Rate, Rate, Probe, Rate, Rate,
No. lb/hr 1b/hr Ratio Ratio [t/sec 1b/Et No. Ib/hr Ib/hr Ib/hr Ib/hr Ib/hr Ib/hr
298 180 4.9 37 86 87 .0531 39,300 2.4i 2.05 0.36 2.46 0.81 3.27
299 87 2.46 0.41
296 87 2.4i 0.36
297 87 2.31 0.26
295 87 2.62 0.57
3lJl 87 .0530 2.65 2.11 0.54
303 87 .0528 2.46 2.30 0.16
302 87 .0530 2.37 2.11 0.26
300 87 2.46 0.36
-------
64
TABLE 15. EXPERU!ENTAL DATA FOR ROUCH-WALL AND HEATED TEST SECTIONS
4 6 7 8 9 10 11 12 13
Hean Orifice Test Section Fuel Rate Chamber No. 1 No. 2 Sample
Drop Air Chamber Air Wall Film To Hixture Sk imme r Skimmer Skimmer Fuel
Run Size, Flow, Vacuum, Pro be Temperature, Tempera ture. F Generator, Flow, Flow, Flow, Cone.
No. microns scfm in. Hg Position F Inlet Out Ie t lb/hr lb/hr Ib/hr lb/hr ug/em3
Rough Wall
312 14 40 8.5 Top 69 67 67 20.6 16.7 0.25 1.62 12.35
311 Center 69 66 66 12.35
313 Bottom 71 67 66 16.0 0.11 1.28 13 .35
Heated Iole t Air
319 14 40 8.5 Top 125 99 95 20.6 32.4
318 Center 120 97 94 13.3 0.08 0.19 31.2
317 Bottom 125 99 93 30.2
Hea ted \,a 11
.
L':! 1[, ~'.. 40 8.5 Top 80 20.6 16.2 20.6
320 Center 80 78 141 0.11 0.07 18.9
321 Bottom 80 77 146 16.4 21.0
,\1] rUII$ \vith 90-degree, 4-inch-bend-radius test section6.
TABLE lb. REDUCED DATA FOR ROUCH-WALL AND HEATED TEST SECTIONS
Test Local Local Average
Section Test Total Entrained Total Wall Total
1\ j r Net Fuel Air Section Fuel Vapor Liquid Fuel Film Sampled
Mass to Tes t Entrained Saturated Velocity, Air Rate to Fuel Fuel Rate to Fuel Fuel
Run Flow, Section, Air/Fuel Air/Fuel average Densit3' Heynolds Probe, Rate, Rate, Probe, Rate, Rate,
No. Ib/br lb/br Ra t io Ratio ft/sec Ibl ft No. Ib/hr lb/hr Ib/hr Ib/hr Ib/hr Ib/hr
Rough Wa 11
312 :1.3 54 94 85 .0540 39,100 1.85 1.91
311 180 97 1. 85 1. 85 1. 90 1.45 3.35
313 4.5 40 94 2.00 1. 85 0.15
Heated Inlet Air
319 49 90 .05l2 37,500 4.85 3.39 1.46
318 180 J.3 25 51 4.67 3.34 1.33 4.68 0.19 4.87
JIJ 49 4.53 3.27 1.26
Heated Wall
n'! 4.4 41 92 (2) .0499 37,500 3.09 0
32 () 18\1 80( I) 2.83 0 3.02 0.07 3.09
32 \ 4.2 43 82 (I) 3.15 0
Notc: All run::; with 90-degrec, 4-inch-bend radius test sections.
(1) At in lcl.
(2) Average.
-------
6S
same for all tests, 20.6 lb/hr. A maj or fraction of the fuel' to the mixture
generator is removed by the chamber skimmer--this is the fuel that is not
entrained by the air in the mixture generator and which impacts on the mixture-
generator-chamber wall and must be removed. A small amount of additional fuel
impacts on the wall of the transition section between the mixture generator
and the test section and is removed by the No.1 skimmer just upstream of
the test-section entrance. The net fuel rate entering the test section should
equal the average entrained fuel rate sampled by the probe at the exit plus
the fuel rate to the No.2 skimmer at the test-section exit.
Mass-balance data are presented in Table 17, from which it is
evident that a good mass balance was not obtained. On the average, there is
a systematic error of 1.4 lb/hr of fuel not accounted for--about 7 percent of
the gross fuel rate to the mixture generator, but a much higher percentage
of the net fuel delivered to the test section.
TABLE 17.
FUEL MASS- BAlANCE SUMMARY
Tes t Sec tion
Air 90-Degree 90-Degree
Flow, Straight 30-Degree Bend, Bend,
cfm Section Bend 4-inch radius O-inch radius
20 Fuel input, 1b/hr 3.1 3.0 3.0 3.0
Fuel collected, lb/hr 1.1 1.2 .w. .L.!
Difference, Ib/hr 2~0 1.8 1.7 1.9
Percent collected 36 41 42 36
40 Fuel input, Ib/hr 3.6 4.6 3.2 4.4
Fuel collected, 1b/hr 1..& ~ 1.& 2.9
Difference, lb/hr 1.0 1.7 0.2 1.5
Percent collected 71 62 9S 66
70 Fuel input, lb/hr 5.2 5.0 4.8 5.8
Fuel collected, 1b/hr l.J! 3.7 3.7 l.J!
Difference, 1b/hr 1.4 1.3 1.1 2.0
Percent collected 73 75 77 65
-------
66
An obvious possible source of error is the result of computing the
net fuel flow to the test section as the relatively small difference between
the fuel to the generator and the fuel removed by the chamber skimmer.
possible errors include the following:
. Nonrepresentative or nonisokinetic sampling of the
Other
entrained fuel
. Overloading the No. Z Skimmer at the test-section exit,
thus not collecting all of the wall-film fuel
. Errors in the determination of the fuel concentration
in the mixture sampled by the probe.
A three-point sampling traverse would not be expected to yield a precise mass
balance; also, no mathematical correction for nonisokinetic sampling conditions
was attempted because of the unreliability of such corrections for pulsating
flow. However, such errors could hardly have been large enough to account for
the mass imbalance.
In examining the data, it is apparent that the values for the total
amount of fuel sampled, as a function of the air-flow rate, are reasonably
consistent, while the data for the fuel removed by the chamber skimmer are not.
It has been concluded, therefore, that the chamber skimmer data are not accurate,
and the cause of t~e erratic data is believed to be carry-over of fuel out of
the trap used to measure the skimmer flow rate.
Consequently, we are inclined
to disregard the computed values for fuel input rate to the test section, and
rely on the observed values of fuel rate at the test-section exit.
Total Sampled Fuel Rate.
It is believed that the No. Z Skimmers
effectively remove all of the fuel flowing as the wall film, and that the simple
volumetric measurement of this fuel rate is relatively accurate--on J:he order
of tz to 5 percent.
The values for the fuel concentration in the samples collected by
the probe are obtained in a less direct manner, and are therefore possibly
subject to greater error. As a check on this, as described previously, samples
of the same concentration were run directly to the hydrocarbon analyzer and
through the probe sampling system. The main probable source of error would
be "hangup" of hydrocarbon within the sampling system, and the checks indicated
that this was not a problem. Values for the total fuel collected by the probes
are estimated to be accurate within :tS percent.
-------
67
Vapor Flow Rate.
The computed vapor flow rate is subject to errors
in measuring the mixture temperature at the test-section exit and is also sub-
ject to the validity of the assumption that the vaporized fuel is in equilibrium
with the liquid fuel. A 2 F error in mixture temperature measurement is equiv-
alent to an error in vapor flow rate of 0.06 lb/hr at 20 cfm, 0.12 lb/hr at 40
cfm, and 0.22 lb/hr'at 70 cfm.
Comparing these possible errors to the observed
flow rates indicated that up to 20 percent error in values for vapor flow and
errors up to 50 percent in entrained liquid fuel flow rate could be anticipated.
Accordingly, values given for entrained fuel should be considered qualitative
at best.
Replication Resu~ts.
The data presented in Tables 13 and 14 were
obtained to provide more precise information on mixture stratification by
traversing with 9 points instead of the usual 3. These data also provide a
check on one data set, the 40-cfm air-flow runs with the 90°-bend, 4-inch-
radius test section presented in Tables 9 and 10.
A comparison of the two
sets of data indicates that fuel-transport behavior is substantially the same
in terms of total fuel entrained, wall fuel-film flow rate, and degree of
entrained-fuel stratification. Only the values for absolute amount of entrained
liquid fuel are substantially different.
Overall Accuracy.
Measurements of total entrained fuel and wall
fuel-film flow rate are believed to be sufficiently accurate to characterize
the fuel transport in the experiments.
Determination of the relative amounts
of entrained fuel occurring as liquid and vapor is subject to substantial errors;
nevertheless, the results are believed to provide a valid qualitative indica-'
tion of fuel-droplet transport characteristics.
Test-Section Flow Conditions
Test-section average air velocities in the experiments conducted
range from about 50 to 115 it/ sec; the resulting Reynolds numbers range from
about 10,000 to 69,000. Thus, air flow is well into the turbulent regime
under all experimental conditions. Since there is a sharp contraction and a
short round-to-square transition section just upstream of the test section,
-------
68
it is unlikely that anything approaching a fully established flow condition
is established at the test-section inlet.
Columns 22 and 24 of Table 4 indicate that there is some stratifi-
cation of the entrained fuel at the test section inlet, particularly at the
higher air velocities.
A higher concentration of fuel in the center of the
section is indicated, which could be caused by the upstream sharp contraction.
The variation in entrained fuel is reasonably symmetrical.
In all cases, the observed test-section wall temperature is a few
degrees colder than the surrounding air temperature. Consequently, the flow
is not adiabatic, although the heat gain should be negligible because the
plastic test-section material is not a good,heat conductor. As discussed
previously, it is believed that the liquid and vapor fuel components are sub-
stantially in equilibrium. Because of the low volatility of the test fluid,
compared to gasoline, air-fuel ratios at saturation range from 80 to 150,
depending upon the fuel temperature.
Droplet Impaction
Figure 23 shows the amount of liquid fuel impacted on the test-
section wall as a function of air-flow rate and test-section geometry. As
would be expected, the least amount of fuel was impacted in the straight
section, and increasingly larger amounts of fuel were impacted in the 3D-degree
and 90-degree bends.
The fact that a greater amount of fuel was impacted in the long-
radius 90-degree bend than in the sharp-cornered 90-degree bend is a surprising
result; however, there is some reason to believe that reentrainment is
responsible for the' apparently lower impaction in the sharp-cornered test
There is additional support for this hypothesis in later discussion.
Figure 24 shows the same data as Figure 23, expressed as a fraction
of the total liquid fuel flow. The lower curve shows a surprisingly large
fraction of the liquid fuel impacted in the straight test section--over 50
section.
percent at 20 cfm, decreasing to 23 percent at 70 cfm. From this, it is
apparent that flow-induced turbulence plays a significant role in droplet
-------
69
/.2
14- micron mean droplet size
Straight section
/.0
0.8
'=
~
,n
~ 0.6
~
r\
Q)
:J
Ii.. 0.4
0.2
0.0
o
10
20
30 40
Air Flow Rate, scfm
50
60
70
FIGURE 23.
WALL FUEL-FILM FLOW RATE AS A FUNCTION OF AIR FLOW
RATE AND TEST-SECTION GEOMETRY
-------
70
1.2
14- micron mean droplet size
1.0 7--\
-0 ..,-goo bend
-------
71
impaction.
The fact that a lower fraction of fuel is impacted in a straight
section as air velocity is increased is not necessarily surprising if an
analogy to heat transfer is made. As air velocity is increased in a duct, the
amount of heat transferred per degree of temperature difference increases, but
less than in proportion to the mass flow of air.
Since heat transfer and
droplet impaction occur by analogous physical mechanisms, one could expect
the fraction of droplets impacted to decrease as air flow increased.
From the curve for the 30-degree bend in Figure 24, it can be
inferred that at the lower air velocities, turbulence is more important than
turning angle in causing droplet deposition.
However, at 70 cfm,the effect
of turning predominates.
The curve for th~ 90-degree, 4-inch-radius bend
indicates that the high-air-flow-rate effect of turning is greater, as would
be expected.
The curve for the 90-degree sharp-cornered bend shows a departure
from the trend of the other three curves, with 100 percent impaction at 20 cfm
and lesser impaction at higher air flow rates.
Again, reentrainment is believed
to be a significant factor in this departure. At flows above 20 cfm, it is
possible that 100 percent droplet impaction is still experienced, but that
reentrainment is negligible at flow rates below 20 cfm.
Figures 25 through 28 show how the entrained fuel droplets are dis-
tributed at the test-section exit, as a function of air-flow rate and test-
section geometry. Figure 25 shows a reasonably flat fue1-droplet-flow distri-
bution at all air flow rates with the straight test section. This could be
expected because of the apparent hig~ degree of turbulent mixing occurring within
the test section. Figure 26 indicates a flat fue1-droplet-f1ow profile at 20
cfm for the 30-degree-bend section, but with increasing stratification toward
the outside of the bend (bottom) at higher air-flow rates.
Figure 27 shows a somewhat increased degree of stratification with
the 90-degree, 4-inch-bend test section at 20 and 40 cfm, as would be expected.
No curve is shown for 70 cfm in Figure 27 since all the droplets were impacted
at that condition.
Figure 28, for the sharp-cornered bend, shows no entrained liquid
fuel at 20 cfm, highly stratified droplets at 40 cfm, and an extreme degree of
stratification at 70 cfm with droplets only along the outer wall of the bend.
-------
c:
.g
'Vi
!r Center
Q)
£j
o
~
Q.
c:
o
-
III
~ Center
Q)
£j
o
~
Q.
BOllomo.o
72
Reor
70 cfm
20 cfm
Front
0.0
1.4
0.4 0.6 0.8 1.0
Locol Entrained liquid Fuel Rate, Ib/hr
0.2
1.2
FIGURE 25. ENTRAINED-FUEL-DROPLET STRATIFICATION AT STRAIGHT
TEST-SECTION EXIT
Top
1.4
20 cfm
0.2
0.4 0.6 0.8 1.0
Local Entrained Liquid Fuel Rate, Ib/hr
FIGURE 26. ENTRAINED-FUEL-DROPLET STRATIFICATION AT 30-DEGREE-
BEND TEST-SECTION EXIT
-------
Top
c
o
....
~ Center
Q.
Q.)
.a
o
...
Q.
Bottom
0.0
Top
c
.Q
-
'Vi
~ Center
Q.)
.Q
o
...
Q.
Bottom
0.0
20
cfm
40 cf m
73
04 06 Q8 1.0
Local Entrained liquid Fuel Rate, Ib/hr
1.2
1.4
FIGURE 27.
ENTRAINED-FUEL-DROPLET STRATIFICATION AT 90-DEGREE.
4-INCH-RADIUS-BEND TEST-SECTION EXIT
0.2
0.4 0.6 0.8 1.0
LDcol Entrained liquid Fuel Rate, Ib/hr
1.2
1.4
FIGURE 28. ENTRAINED-FUEL-DROPLET STRATIFICATION AT 90-DEGREE,
O-INCH-RADIUS-BEND TEST-SECTION EXIT
-------
74
This figure does not strongly support the hypothesis that reentrainment is a
significant factor, nor does it particularly detract from it.
From the droplet-impaction experiments, it is clear that ultrafine
atomization does not make it possible to avoid fuel-droplet impaction because
of the role of flow-induced turbulence in causing impaction. Nevertheless,
it appeared possible to avoid impacting at least half the fuel under the range
of air-flow rates investigated with a l4-micron mean-droplet size in both
straight and 30-degree-bend test sections. The effects of bend turning angle
are more significant at higher velocities and predominate over turbulence-
caused effects at 70 cfm.
Over half the fuel was impacted in the, 90-degree-bend
test sections under all conditions with l4-micron mean-droplet size.
The fact that the sharp-cornered 90-degree bend resulted in less net
fuel impaction, possibly due to reentrainment, should not necessarily be used
to indicate that the sharp-bend configuration is superior to the long-radius
bend for intake manifolds. In this case, the reentrainment is probably the
result of greater liquid-fuel hold-up in the bend--a condition leading to time
variations in mixture ratio and poor response to transients.
Runs 311-313 were run with a rough-walled test section (90-degree,
4-inch-radius bend) to check for possible effects of wall surface on droplet
impaction. Data for these runs are presented in Tables 15 and 16, and can be
compared to Runs 229-231, in Tables 9 and 10 for equivalent test conditions
for a smooth-walled section.
This comparison shows about twice the fuel
skimmed off the wall with the rough surface-- a surprising result in that no
difference in the rate of droplet impaction was anticipated.
No explanation
for this result is available, nor is the significance of the result apparent
at this time, except for one possibility.
There is the possibility that
reentrainment is significant in all the tests conducted on droplet impactton,
and the rough wall inhibits reentrainment. It should be emphasized that ~here
is no evidence that this is the case. However, if it were true, smooth-w~lled
intake manifolds could be expected to produce superior engine response
characteristics.
Generally, then, it can be concluded that ultrafine atomization,
minimum manifold-passage turning angle, long bend radii, and low air velocity
are all conducive to low droplet impaction; however, some impaction will occur
under any circumstances if there are fuel droplets in the air stream.
It is
-------
75
therefore further concluded that nearly complete vaporization in the mixture
generator of auto-engine induction systems is highly desirable.
Wall Fuel-Film Transport
Wall fuel-film transport observations were made purely on a visual
basis, and the information so obtained relates mainly to flow patterns and areas
of liquid holdup. Figures 29, 30, and 31 are sketches illustrating the observed
flow patterns. The patterns for the 30-degree bend and the long-radius 90-
degree bend were similar and did not vary appreciably with air-flow rate.
Liquid deposition on the walls was apparently uniform upstream of the bend; at
the bend, liquid was observed to move across the side walls from the outer wall
to the inner wall. It is highly likely that this flow pattern is the result of
secondary air flows along the side walls of the bend which are caused by the
higher static air pressure along the outside wall of the bend.
Liquid swept toward the inside of the bend accumulates at the
center of the inside wall, where the liquid film is noticeably thicker. The
increased droplet deposition along the outer wall is apparently well distributed.
Figure 31 shows the wall fuel-film patterns observed in the sharp-
cornered 90-degree bend. With this geometry, there are pockets of liquid-fuel
holdup on both inner and outer walls. Intuitively, it would seem that such a
flow condition is highly conducive to reentrainment.
These observations were made with transparent, smooth-walled test
sections, while intake manifolds are typically rough surfaced sand castings.
We believe that the general flow patterns will be similar for both smooth and
rough-walled passages; however, film thicknesses and velocities would be
altered.
Also, the possibility for rentrainment might be different with a
rough surface.
No conclusions have been drawn in this study regarding absolute
film thickness and velocity.
The character of the fuel-film surface on smooth-walled test sections
in most cases would be described as "rippled", but not thick enough to become
"wavy".
-------
76
.
..
Top View
Side View
FIGURE 29.
WALL FUEL-FILM FLOW PATTERN IN 3D-DEGREE
TEST SECT ION
-------
77
~
//
.
..
Top View
Side View
.
..
Bottom View
....
FIGURE 30. WALL FUEL-FILM FLOW PATTERN IN gO-DEGREE,
4-INCH RADIUS TEST SECTION
-------
78
~
/
~~~
---~
(
Accumulation
of liquid
...
-
Top View
.-
Accumulation
of liquid
..
..
Side View
...
...
Bottom View
...
FIGURE 31. WALL FUEL- FILM FLOW PATTERN IN 90- DEGREE,
O-INCH RADIUS TEST SECTION
-------
r
I
79
V a ~~x.i~.!. iOll
The results of the experimental work provided information on
vaporization both from liquid-fuel droplets and from the liquid wall film.
For the case of droplet vaporization) this information is largely qualitative)
since rate of vaporization could not be measured directly in the experimental
apparatus.
As far as the authors could determine) existing literature infor-
mation on droplet and film vaporization cannot be applied to compute rates of
vaporization that might be occurring in the experiments conducted.
Dr~et Vaporization.
In conducting experiments on droplet dynamics)
it was desirable to have test conditions that would result in a constant droplet
size. This could be achieved by two possible circumstances: (1) droplet vapori-
zation in the mixture generator so rapid that equilibrium conditions are estab-
lighed in the test section) or (2) droplet vaporization so slow that the amount
of vaporization in the test section is negligible.
As has been discussed pre-
viously in this report) the former alternative prevailed and necessitated the use
of a test fuel of low volatility in order to establish droplets in the test
section.
Accordingly) we know droplet vaporization was rapid; unfortunately)
we cannot characterize accurately how rapid. However) the calculated residence
time of the air in the mixture generator provides a lower bound on the vapori-
zation rate.
At 70 cfm air rate) this residence time is approximately 0.24
second.
Mixture residence times in current conventional induction systems are
on the order of 10 to 100 milliseconds; consequently) the ,residence time in
.the mixture generator of the experimental apparatus represents from 2-1/2 to
25 times the residence time in a typical induction system.
The data for Runs No. 317-319 in Tables 15 and 16 were taken with
air entering the mixture generator heated to 125 F.
Data for the same test
section without preheated air are given for Runs 229-231 in Tables 9 and 10.
A comparison of these data shows that about 60 percent more fuel was captured
by the air in the mixture generator with the heated air.
If aerodynamic
entrainment was the principal mechanism by which the air captures the fuel)
the effect of inlet air temperature on the amount captured should be slight)
and the main effect of temperature would be to vaporize a greater fraction of
the entrained fuel.
-------
80
The fact that more fuel is captured with the heated air indicates
that vaporization during the aerodynamic entrainment is appreciable. The
zone in which entrainment occurs is of small volume relative to the overall
volume of the mixture generator, consequently, it can be inferred that sub-
stantial vaporization occurs in a fraction of the total residence time of the
air in the mixture generator.
Therefore, although quantitative data on rate of vaporization were
not obtained, there is reason to expect that fuel droplets in the vicinity of
14 microns size would essentially approach equilibrium vaporization in air in
milliseconds, and could vaporize appreciably in times as short as 10 milliseconds.
Heated-Wall Tests.
Data for tests at 40 cfm with a heated-wall
test section (90-degree, 4-inch-radius bend) are given .in Tables 15 and 16,
Runs 320-322.
Wall-temperature data are presented in Table 18.
Compared to
Runs 229-231 in Tables 9 and 10 for the same conditions with an unheated test
section, it is seen that the amount of wall-film fuel captured by the No.2
skimmer is 0.07 lb/hr with the heated section and 0.76 lb/hr without heat.
Thus, it is indicated that the effect of heating the test-section wall to
190 F is to vaporize about 90 percent of the fuel that would otherwise impact
and remain on the test-section wall.
Another effect is that the mixture temperature rises about 64 F in
the test section; thus, the mixture, which was saturated with fuel vapor at
the test-section inlet, is not saturated at the exit.
This leaves us with no
information on whether there is entrained liquid fuel at the exit under these
conditions.
by two
before
Heating the test-section walls can reduce the fuel-film flow rate
possible mechanisms: (1) by causing the entrained droplets to vaporize
they impact, or (2) by vaporizing the wall-film fuel.
The wall-temperature data presented in Table 18 show that the outside-
of-bend wall temperatures downstream of the bend are a few degrees lower than
the average, indicating that impacted fuel droplets are probably cooling the
wall. Consequently, it seems unlikely that all or most fuel droplets are being
vaporized before they impact. The residence time of the mixture in the test
section at 40 cfm air input is about 16 milliseconds, and it is likely that
-------
, .
',',
1,- ','
It. '.';
. ;, .! ' ,
al
TABLE 18.
WALL-TEMPERATURE DATA FOR
HEA TED TEST SECTION
Thermocouple Thermocouple Temperature,
Location Number F
Inside wa 11 2 202
" 3 191
Side wall 6 196
" 1 187
II 8 186
II 9 187
Outside wall 4 194
" 5 198
II 7 180
II 10 180
Thermocouple locations are shown in Figure 10.
14-micron fuel-droplet vaporization is not complete within that duration.
There is no good measure of whether droplet vaporization is appreciable in
the test section with the heated wall.
The amount of heat transferred from the wall in vaporizing the fuel
is small compared to the sensible heat transferred to the air. Assuming 0.7
lb/hr of fuel is vaporized and that the latent heat of vaporization is 150
Btu/lb, 105 Btu/hr are transferred to vaporizing the wall fuel-film. Estima-
ting that the air is heated 64 F in the test section, a heat input to the air of
2770 Btu/hr is computed.
Although the mechanisms by which the fuel is vaporized in the heated
test section is not clear, it is clearly significant that the wall fuel-film
was virtually eliminated by the effect of the 190 F wall temperature.
-------
82
Overall Conclusions
-- --- _.-
It is apparent that, with fine atomization, minimum manifold-
passage turning angle, long bend radii, and low air velocities, fuel-droplet
impaction can be kept ~ow, but it cannot be avoided.
Fortunately, ultra-fine
atomization is also highly conducive to rapid droplet vaporization, and
.'. .. ", .
indications are that droplet vaporization with droplet sizes below 20 microns
occurs rapidly enough to achieve near-equilibrium vapor concentration in a
practical mixture generator at moderate engine speeds, and possibly to achieve
appreciable vaporization at higher engine speeds.
Even though some droplet impaction is apparently unavoidable, it
is evident that, with fine fuel atomization, the wall fuel-film can be vir-
tually avoided by the use of moderately heated intake-manifold passages.
-------
83
DESIGN CONCEPTS
The original objective of this part of the study was to design,
construct, and demonstrate a laboratory prototype fua1-induction system
having the capability of improved fuel-air mixing and distribution.
This
system was to include a mixture generator, a vaporization section, and an
intake manifold.
While the prototype fuel-induction
system was intended for
initial use as a laboratory apparatus, one of the design constraints was that
the mixture-generator and intake-manifold should be within size limitations
that would ultimately permit use in the space available in standard automo-
biles.
The vaporization system, on the other hand, was considered for use
only as a laboratory tool for providing a fully vaporized-fue1-air mixture to
an engine.
As previously mentioned, the scope of this part of the study was
amended to include conceptual design studies; and the detail design, fabrica-
tion, and demonstration tasks were deleted.
Induction System Concept
General Description
Figures 32 and 33 are layout sketches of the proposed fuel atomiza-
tion system and intake manifold. The system consists of an air inlet and filter
housing, a fuel atomizer, a mixing chamber, a throttle, and a manifold. The
air cleaner diameter and its height above the engine have been made approximately
the same as for conventional induction systems so that this system would fit
under the hood of standard automobiles. The air cleaner is also designed to
house a standard filter element.
The configuration shown in Figures 32 and 33 will fit a Ford 351
crD V-8 engine. This engine is available in both automotive and industria1/
marine versions, and was se1ect~d for the primary application because the
inlet ports on the cylinder heads are spaced uniformly. The uniform inlet-port
spacing simplifies the layout of intake passages around the circuiar manifold
plenum.
-------
I
13.5
Fuel-
-Atomizino oir
Hortmann-whistle
air-atomizino nozzle
o <:::>
~ .
I,
~
r
~
.,
"
i
!
2.5
, ,
I
I
/
II
'-..1
\~
I
3.6
6.75
lG'25 Throttle cylinder
I Width of port is
2x height here
./~/
. '-"
.'
/
/
/'
/
\
Section A-A
FIGURE 32. SKETCH LAYOUT OF PROTOTYPE IMPROVED INDUCTION SYSTEM -SECTION VIEW
Throttle coble, one of three
equally spaced around,
throttle cylinder '
00
~
. . />0 . ~ / Cylinder head
'v' ~
?
-------
---
.......,,---
------ //
-, /
...... ..... .." .- "'''--- /
" ...-~-- '- ./
,,"'ttt:' - /'.
/ .... - / '. - - --
--->~~ "\ //'/~<~n A
/' /-/ \\, \\,i/ /', '-', "" y
.' I ' \ 1 I - --, / '
/ / \ \, 1 )_-l\i
I \ 1,/ I - - I I '
I ./ rr'-' ... IiI;, -- II I
/ I, - - -1- \ - I / J',
i 'I I - - - ' ." -,.,'.,~, I / ;' ---- I , "
I " I - - \ -- .' 1/ """ I, .
/ I --, / 1''/ I I
I :! : ';y/ 2.0 (typ) / / / /--:-::J.~
'I.' ~.<
" L_- I /' /
~'''---')--''-, , / /1
A "I / /
Lh! ""-,J "" <-"':"
I " I ,
\ / \ "
/ , / / '''''''- 'i",
\ \ ,'r' -;...,;---/-' /'/\ -:: i
: ; ,- , , ' " II I
\-111 /'" ,-- I"
'\ L4 -"'" /1 \ - - I : I
:1 - /",,/1 \ ---,-__11
\ ~-:_;----/ / ,/\ \ - ,,- /
''\ " I I, \ ./ ,/
'\ '\. / / \ ~\
"", I " \ \
". "" / 1 \ \ "
I 1 " "
/ I ,
"I ,
...,,,,,... I
,," "', I "
---- ""'..,' ~::----
-~/
85
. .. ,.---...
16.~
,
I
~
.~ ._--..~-_._'- -.-.. ,.,___n_.
"
10.85
FIGURE 33. SKETCH LAYOUT OF PROTOTYPE IMPROVED INDUCTION SYSTEM-PLAN VIEW
I
I
I
I
I
I
I
I
/
I
i
4.5 I typ)
-------
86
Figures 34 and 35 are sketches of a proposed intake-passage
layout for an engine with paired inlet ports
as the Chrysler 318 crD and 340 crD engines.
two ports on each side are more tortuous than
in the cylinder head, such
The passages to the middle
those for the spaced-port
manifold previously illustrated. However, the bends are not quite as sharp
as they appear to be in Figure 35 because there is an elevation change
of about 1-3/4 inches from the manifold plenum to the cylinder head port.
The dimensions given in these and subsequent figures in this
section of the report are the more important or critical dimensions.
Dimensions not given are generally arbitrary.
Width of port is
2x height here
Horizontal run here ~'....
"
Throttle cylinder
"
Section A-A
FIGURE 34. SKETCH LAYOUT OF INTAKE-PASSAGE CONFIGURATION FOR PAIRED-PORT
ENGINE - SECTION VIEW
-------
87
... 9.55 -1
I
I
,
I
12.6
"
A" .
t_.-
-,
,
"
" ,
, ,
\ ,
, ,
\ \
\ \
II
\ \
\ \
"
,
"
I
,
I
,
I
\ \
\ \
~
~
\
J
I
\
\.
"
,
..... ;;.
/ /
I ./
I ./
/-
,/
r
./
I
I
~
--- -" /
... -"
/
/f
'i
1
1
1 1
/ /
" ,
- -
.-
...
....
..-
I ....
, - ~/
, I - - - -:-'-
'.'
f
.. ..)
,
,
"
,
\
\
\
/
/
/
I
\ 1
V
1- .. .. '.' '!
- - t - - -I I ,
....""" I,
/
/
i . ..~. ... t' . ....
! I
- - - - !
.... - - - -1A
,/ ..,-
/,/ - , :~
1/ I--n'
/ / I
, / ....: ! I ,
II I ~ :.. .J'.J
I, ,I
II I I
II \
II I I
I I
j
./
or'-''''-
....-/
1\
1 \
/
/
/
Plan View
..'
...
...
-"
1.1
..j
"
.-
/
.'
./
I
I
1
. I
1
-
./
....-
/
,/
-"
.-
....
<.,
7.95
....
,
'\,
\
\
I
I I
" / I'
\,
\, I I
\'
II !
o I
\ 1\ , I
\ \\ _....~,
,\ '...... I I I
\ , I I I ..
\ " \ \,' " -1 ~ ~.._.._.l-
, ' " I I
, ...... I
'" .,'---+111
" --~ I
, I II
',.. I II
'" - I I J
--r--..II
~---~
FIGURE 35. SKETCH LAYOUT OF INTAKE-PASSAGE CONFIGURATION FOR PAIRED-PORT
ENGINE - PLAN VIEW
-------
88
Fuel Atomizer
.-.----- --
Several alternative atomization systems appeared promising for
the automotive application.
These were: the piezoelectric ultrasonic atomizer,
the impinging-jet air atomizer, the Hartmann-whistle-type atomizer, and the
spinning-disk atomizer.
In the ultrasonic atomizer, liquid fuel flows over a vibrating
surface and is atomized by forces generated in the fluid by mechanical
agitation.
The piezoelectric ultrasonic atomizer consists of an electronic
power supply which drives a piezoelectric disk.
Attached to the disk is a
cylindrical horn with a sharp outer edge which vibrates in a radial mode.
Fuel flowing down the outer surface of the horn is atomized off the sharp edge.
The mean droplet size of liquid atomized ultrasonically is dependent
on the frequency of vibration.
A practical frequency range for units intended
to deliver substantial flow rates is 25 to 100 KHz. An atomizer operating at
100 KHz can deliver droplet sizes in the range of 10 to 20 microns(]).
The power requirement for ultrasonic atomization varies from about
5 watts for 2.1 Ib/hr of fuel oil (8) to about 23 watts for 160 lb/hr of water(9).
In the impinging-jet air atomizer the energy of compressed air is
used to atomize the fuel. In a simple configuration, a high-speed air flow
is directed around the outside and approximately perpendicular to a jet of fuel.
Small droplet sizes can be obtained with impinging-jet atomization
because the energy contained in the air stream can be independent of the
quantity of fuel being atomized; that is, a large high-velocity air jet can
be used to atomize a small amount of fuel.
The power requirement for impinging-
jet atomization is essentially the pumping work of the air.
The Hartmann-whistle-type atomizer involves a jet of high-velocity
air which is impinged upon the open end of a small cavity(lO). This jet
whistles with such intensity as to provide strong local shock waves in the
space between the air nozzle and the cavity. Fuel introduced around the
periphery of the air jet is atomized in the highly turbulent flow pattern
around this cavity.
-------
89
Droplet size capability of the Hartmann-whistle-type atomizer is
of the same order as the piezoelectric ultrasonic atomizer.
Air consumption
is generally higher in this type of atomizer than in the impinging-jet
atomizer, but smaller droplet sizes are attainable.
~inning-disk atomization is based on the centrifugal acceleration
of the fuel to a high velocity and subsequent discharge into the air stream.
The fuel is introduced at the center of the disk, flows radially across the
surface of the disk under the action of centrifugal force, and leaves the
disk in a thin liquid sheet which immediately breaks up into uniformly sized
droplets. Droplet size is dependent on disk radius and disk speed. Power
requirements are dependent on fuel feed rate, disk speed, and disk radius.
Spinning-disk atomizers are capable of producing droplet sizes
well under 20 microns, and energy requirements are in the 400 to 500 watt
range for maximum fuel flow.
Each of the promising alternative fuel atomization systems was
evaluated on the basis of simplicity, compactness, energy requirements, and
minimum-droplet-size capabilities. Although a desirable droplet size of
under 20 microns has been indicated by both theoretical considerations and
experimental data, it has been stipulated that the fuel atomizer to be pro-
posed as a prototype should be capable of producing a range of droplet sizes.
With this capability, the propotype can be used to investigate the effects of
droplet size on performance in an actual engine.
Table 19 summarizes the results of the fuel-atomizer evaluation.
The Hartmann-whistle-type atomizer appeared to be the most promising, as it
is fairly well developed in this application, is simple and compact, and can
be designed for a wide range of droplet sizes and fuel flow rates.
The actual fuel nozzle selected for the prototype induction system
is similar to a commercially available unit which is being used in a modified
form in a low-emission-burner development program at Battelle.
The nozzle
operates on the Hartmann whistle principle, with 'an air and fuel jet directed
at sonic velocity against a cup-like deflector.
of finely atomized fuel results.
A cone-shaped spray pattern
Figure 36 is a sketch of the proposed nozzle showing the spray
pattern.
-------
90
TABLE 19.
COMPARISON OF ALTERNATIVE ATOMIZATION SYSTEMS
Piezo-E1ectric
Ultrasonic Impinging Jet Hartmann Whistle Spinning Disk
Lowest practically attainable
mean drop size 20~ <20~ «20~ 20~
Fuel flow rate capacity Adequate with possible Adequate Adequate Adequate if stacked
limitations disks used
Energy required for maximum
fuel flow rate of 36 1b/hr -15 watts .-.500 watts --450 watts .-.500 watts
Components required Piezo crystal Nozzle Nozzle Disk (s)
Power supply Air pump Air pump High-speed motor
Fuel control Fuel control Fuel control Fuel control
Space requirements Medium Sma 11 Sma 11 Large
No 69 drill through --,
perpendicular to centerline (8 holes) ,
Fuel spray pattern
0.4
, ,
"
"
" '
", "
..
"
," "
"
Resonator cup
Wire suppc;>rt
..
,"
, , ,
"
, ,
FIGURE 36. SKETCH OF PROPOSED AIR-ATOMIZING,
HARTMANN-WHISTLE-TYPE FUEL NOZZLE
..
-------
91
According to data supplied by the nozzle manufacturer(lO) and to
Battelle's experience in the low-emission-burner development program,
an atomizing air supply of about 7.1 scfm at 22 psig is required for the full
mixture flow rate of the engine (36 lb/hr). At this condition the air pump
power requirement would be about 0.6 hp. The fuel droplet size in a
Hartmann-whistle-type fuel atomizer is a function primarily of the ratio of
atomizing-air flow rate to fuel flow rate.
Larger ratios produce smaller
droplets. There are no reliable data available concerning the specific
droplet size range capability of the nozzle selected. However, the manufacturer
has reported achieving 50-micron-size droplets with water at low levels of
pressure and flow, and droplets considerably smaller than 20 microns at high
pressures and flow rates with water. Given the differences in physical charac-
teristics between water and gasoline, it is reasonable to assume that the
Hartmann-whistle-type air atomizing nozzle illustrated in Figure 36 can be
operated to produce a droplet size range of 15 to 100 microns. The dimensions
given on the sketch have been selected with this objective in ,mind.
Besides the nozzle, this atomization system requires an air supply
and a fuel supply. The atomizing air can be supplied by an air pump similar
to those used for AIR (air injection reactor) emission control systems. This
air pump is belt driven from the engine, and delivers about 10 scfm at idle
and 50 scfm at 100 mph. Corresponding delivery pressures are 1/2 psi at idle
and 15 psi at 100 mph. Fuel can be supplied to the proposed atomizer using
a regular automotive diaphragm-type fuel pump.
Mixing and Vaporization Chamber
The purpose of the mixing chamber is to promote fuel vaporization
and good mixing between the air and the fuel while minimizing the deposition
of fuel droplets on the walls. As with the fuel atomizer , space limitations
are important.
In the mixing chamber configuration selected for the prototype
induction system, the inlet air is introduced through holes in the chamber
walls from an annulus passage outside the chamber.
The hole sizes and pattern
illustrated are intended to produce a symmetrical non-swirling turbulence
-------
92
pattern.
Air flow through large holes at the very top would penetrate to the
center of the chamber and mix with a portion of the fuel spray.
Fifteen
0.45-inch diameter holes would be required in this row.
The row of small holes
near the top should produce low-momentum jets which would be deflected down
along the chamber wall.
Thirty-two 0.'30-inch diameter holes would be
required in this row.
The other two rows of holes pTovide additional air for
mixing with the air and fuel mixture in the region between the center ~nd the
walls of the chamber. Twenty-one 0.30-inch diameter and seventeen 0.45-inch
diameter holes would be required in these two rows. The pressure drop across
the mixing chamber at full flow will be about 4 inches of wa'ter.
It should be emphasized that the optimum mixing chamber configura-
tion can only be arrived at by trial and error. The configuration shown is
considered a good starting point.
The volume of the mixing chamber as described in Figure 32 is
about 0.046 cu ft, which will result in mixture residence times of about
0.14 sec at idle and 0.009 seconds at maximum speed.
This range of residence
times can be expected to produce a high degree of fuel vaporization, at least
at low engine speeds.
Throttle
The throttle as shown in Figure 32 is a cylindrical "sleeve"
which controls the openings between the inlet manifold plenum and the manifold
passages. The throttle sleeve moves vertically to increase or decrease the
openings through which the fuel and air mixture must pass to reach the engine
cylinders. The movement can be effected by cables as shown (3 would be
used to avoid binding) or by pins and a cam or levers from below. When
fully closed, V-shaped slots in the sleeve, one located at each manifold
passage, would allow the fuel and air mixture at idle to pass through.
Figure 37 illustrates an alternative to the sleeve throttle
shown in Figure 32.
In this alternative, the central section of the mani-
fold, or "plug" is moved up and down to control the mixture flow. The
movable element could be centrally supported and actuated from below.
-------
93
,.,L:
-..
Throttle
Spring return to
idle position
FIGURE 37. SKETCH OF PROPOSED INDUCTION SYSTEM WITH PLUG THROTTLE
The relative merits of the sleeve and plug throttle configura-
tions are difficult to assess without conducting further studies. In both
cases at part throttle opening there is a potential for fuel impacting on
surfaces of the throttle and then becoming reentrained in the manifold
passage in an unstable manner. Furthermore, in both configurations the part
throttle position will result in downstream eddies (in the manifold passages)
which may have some detrimental influence on time-based cy1inder-to-cy1inder
distribution. However, either configuration should be a significant improve-
ment over the conventional butterfly throttle plate. Mechanical design will
also play an important role in selection of a suitable mixture throttle.
-------
94
It must be acknowledged that the sleeve valve presents potential
difficulties in both design and manufacturing.
The actuating mechanisms
shown cannot be considered practical designs, and the required tolerances
for the sleeve valve and its se~t and guide could involve considerable manu-
facturing expense.
The plug valve presents similar problems.
An alternative to sleeve-valve or plug-valve throttling schemes
shown in Figures 32 and 37 is to move the throttle to the entrance of the
mixture generator.
In this location, the valve imposes no fuel-impaction
,problems, and, therefore, a conventional butterfly-type throttle could be
used. This approach necessitates designing the mixture generator as a
pressure vessel. It also introduces a minor complication in the design of
the fuel-metering system in that the system must discharge into a variable-
pressure zone instead of a constant-pressure zone.
Figures 38 and 39 are layout sketches of the prototype induction
system showing a suggested configuration for inlet-air throttling. An inlet-
air chamber with IO-gage (approximately 1/8-inch thick) walls surrounds the
mixing and vaporization chamber, and both are bolted directly to the intake
manifold. A circular-cross-section throttle passage is welded to the side of
the inlet chamber.
A baffle at the inlet-air chamber entrance distributes
the air more uniformly around the mixing chamber. The air filter housing fits
essentially air-tight at the throttle-passage entrance by overlapping edges.
A standard air filter is used.
Inlet J1anifold
Design of the inlet manifold for low impaction was based on reducing
the number of 90-degree bends in each passage and providing minimum turning
angle consistent with available space in the remaining bends. A single-plane
central-plenum configuration offered the best approach to achieve these
objectives. With the single-plane plenum, individual manifold passages can be
made fairly short, providing the diameter of the central plenum can be made
large. A large-diameter central plenum is compatible with the design of the
mixing and vaporization chamber and with the use of an air atomizing fuel
nozzle.
-------
95
Air filter
nozzle
Air filter housing-
, I
\.1
II
~ Inlet-air throttle
FIGURE 38. SKETCH LAYOUT OF PROPOSED INDUCTION SYSTEM WITH
ALTERNATIVE INLET-AIR THROTTLE - SECTION VIEW
The inlet manifold configurations shown in Figures 33 and 35 con-
sist of a central cylindrical plenum, equally spaced ports around the circum-
ference of the plenum, a~d 'passages of varying lengths between these ports
and the engine cylinder-head ports. The manifold passages were designed with
bend radii (measured to the outside passage surface) no less than four times
the passage width. At the central plenum the passage ports are about 1 inch
high and 2 inches wide. At the cylinder head the passage ports are approxi-
mately square to match the cylinder-head ports.
The conical hump in the center of the manifold plenum is designed
to provide guidance to the mixture flow entering the manifold passages through
the sleeve throttle opening. The optimum height and shape of this hump would
be experimentally determined.
-------
96
:..--- .----
-- -
.....- --
./'" .........
/ ""
./ ../" --- ---.............. "
/ >/ ~, ~
// " \\
/' / \ \
I / \ \
/ I \ \
I I \ \
( ( I 1
\ \ I I
\ \ / /
\ \ / /
\ " / /
" " / /
" / /
"" ~-- -----/ /
" /"
" /"
........... . ~
--------
FIGURE 39. SKETCH LAYOUT OF PROPOSED INDUCTION SYSTEM
WITH ALTERNATI VE INLET -AIR THROTTLE - PLAN VIEW
-------
97
Idle Capability
With systems using the sleeve- or plug-valve throttle, the idle
fuel-and-air mixture flow can be provided for by notching the throttle sleeve
(or throttle plug) at each manifold-passage pott. The fuel atomizing nozzle
and mixing chamber should be able to provide a homogeneous mixture of finely
atomized fuel and air even under idle conditions because the fuel droplet size
depends only on the ratio of atomizing air flow to fuel flow and not on total
air flow. By proper design of the atomizing air supply system, this ratio can
be maintained high enough at idle to achieve the fine atomization necessary for
good mixing and low impact losses. Thus, it is anticipated that a separate
idle fuel system would not be necessary.
Ac~~lera~ion Enrichment
The air atomizing or Hartmann-whistle-type nozzle does not lend
itself well to acceleration enrichment. Thus, if acceleration enrichment is
required, it will be necessary with this induction system concept to provide for
this enrichment separately. One method of doing this would be simply to provide
an accelerator pump, as most carburetors have, to inject an extra amount of
fuel into the mixing chamber when the accelerator is depressed. The quantity
of fuel injected in this manner should be proportional to the speed of movement
and distance of travel of the accelerator. This accelerator pump could be
designed so that gradual movement of the accelerator did not cause an injection
of excess fuel.
The fuel could be either injected from a point near the atomizing
nozzle or from a nozzle in the center of the hump in the manifold plenum.
The
latter injection point might provide the quickest response characteristics.
Specific configurations for control of idle operation and accelera-
tion enrichment have not been provided in this study. The aim of the proposed
prototype fuel induction system is to demonstrate the potential of improved
fuel atomization, improved mixing and distribution, and reduced droplet impac-
tion for reducing exhaust emissions, particularly NO. If this potential is
x
successfully demonstrated, then the idle and acceleration enrichment require-
ments would be tackled next.
-------
98
fuel Me~ering System
In the proposed induction system, fuel would be supplied to the
atomizer nozzle by a fuel pump. The pressure generated by the fuel pump at
any engine speed would have to be in excess of that required for the maximum
fuel flow rate anticipated for that speed. Control of the actual fuel flow
by a throttling valve would provide the fuel metering.
To insure precise air/fuel ratio control at any speed and load
condition a means would be required to sense air flow.
One such method would
be to provide a venturi in the air inlet and to use the venturi pressure
differeqtial as an air flow signal to control the fuel metering.
Another
approach would be to measure engine speed, intake manifold air pressure, and
air temperature, and to electrically combine these signals in conjunction with
a predetermined engine volumetric efficiency curve.
signal would set the fuel rate.
The resulting output
The fuel metering control would also have to be designed to provide
enrichment for cold-weather starting -- equivalent to the effect of the choke
in conventional carburetors.
Fuel Vaporization
Eftect of ,Fuel Vaporization
on Maximum-Power Capability
Vaporization of the fuel potentially results in a decrease in the
partial pressure of the air in the fuel-air mixture, which, in turn, can
result in reduced engine air breathing and reduced maximum-power capability.
However, a cursory analysis indicates that this will not be a problem and
that a power increase is a more likely result.
If it is assumed that the fuel can be represented by n-heptane,
and if it is further assumed that the mixture is at 100 F and one atmosphere
absolute pressure, the volume of 1 lb of air is 14.1 ft3 while the volume of
0.067 lb of fuel vapor (stoichiometric ratio) is 0.27 ft3 (assuming perfect-
gas-law relationships are valid). The additional volume of the fuel is
-------
99
equivalent to only 2 percent of the air volume; therefore a 2-percent power
loss is the most that could be anticipated.
However, if the cooling effect of fuel vaporization is considered,
an increase in the air density of the fuel-air mixture can be anticipated.
Assuming the average heat of vaporization to be 150 Btu/lb, and assuming that
both air and liquid fuel are initially at the same temperature, a 37.5 F drop
in mixture-temperature can be expected as a result of the fuel vaporization.
If the air and fuel are initially at 100 F, the resulting drop is about 6
percent of the absolute temperature. The net result of vaporization consid-
ering effects on both mixture temperature and air partial pressure is an
anticipated 4 percent increase in engine air flow, assuming an adiabatic intake
manifold. Since the manifold is not adiabatic, something less than a 4-percent
increase would be anticipated in a real engine.
~~t-Spot Vaporization in Manifold
Although ultra-fine atomization of the fuel should preclude the
necessity for a high-temperature manifold hot spot, there may still be s'ome
need for providing moderate, uniform heat in the manifold for vaporizing any
fuel which does deposit out of the mixture. The cone-shaped hump in the
center of the proposed manifold plenum offers a suitable surface for manifold
heating. The interior of this cone could be hollow with passages to the
cavity through the intake manifold mating with exhaust passage ports on the
cylinder heads.
In the case of the alternative "plug" throttle with the movable
cone, it would be more difficult but not impossible to provide the cavity and
passages for exhaust gas flow.
Fuel Vaporization Chamber Concept
The proposed induction system concept, with or without manifold
heating, may under some conditions still deliver a mixture containing fuel
droplets to the engine. Thus, it is expected that geometric and time-based
distribution of the mixture will not be perfect. To provide a means by which
-------
r-~'
100
to evaluate the improvement in distribution that can be achieved with complete
vaporization of the fuel, and to obtain a measure of the amount of heat
required to accomplish this, a separate vaporization chamber was designed
for use as a laboratory device.
The design criterion for this vaporization chamber was that it
have the capacity for completely vaporizing the fuel at maximum air/fuel
mixture flow rate \"hen used either in conjunction with the proposed induction
system prototype or with conventional induction system components.
After consideration of several alternatives, the approach selected
to vaporize the fuel in an induction system was to seek a design in which a
majority of the fuel droplets could be caused to impact on heated surfaces
regardless of the mixture flow rate. Two important additional criteria in
the design sought were compactness and low pressure loss. No attempt was made
to design a system what could fit under the hood of a car.
Several approaches to the design of the vaporization chamber were
investigated before arriving at the final design. For instance, fuel could be
vaporized in a length of heated pipe. From previous experience with heating
air in such an arrangement, it was found that a length of about 20 feet would
be required to heat the air from 70 to 140 F.
With such an arrangement, the
air should be preheated before entering the vaporization section in order to
minimize the length of pipe required.
One means of increasing the contact time of the air with the hot
walls of a heated-pipe vaporization chamber would be to induce swirl with a
twisted tape.
The tape would perform a second beneficial function of centri-
fuging large fuel droplets out to the heated walls. However, a mixing section
would have to be provided to break up the overly rich mixture at the walls
caused by vaporization of the large droplets. Large droplets would probably
impinge on the cool twisted tape at high flows. The resulting fuel film on
the tape would probably be reentrained by the air in an irregular manner.
Vaporization effectiveness of this arrangement would be a function of the mix-
ture flow rate. At low flows, large droplets may not impact on the heated walls.
It was felt that a better approach would be to look for a design
in which a majority of the fuel droplets would be likely to impact on heated
surfaces regardless of the mixture flow rate.
The first impaction-type
-------
101
vaporization chamber investigaed had Thermek tubes.
Thermek tubes are
produced by a process whereby the surface of a tube is mechanically lifted
to form spiraled spines completely surrounding the tube. Fuel droplets intro-
duced perpendicular to a bank of these tubes would surely impact. However,
insufficient heat-transfer information was available on this type of extended
surface ~or predicting air-heating rates.
A number of standard-type heating coils were then investigated.
A heat exchanger configuration was found which combined compactness with high
surface area.
This heat exchanger is a Mcquay Nodel SBB hot water booster
coil with closely spaced fins that are corrugated to produce a tortuous flow
path.
Figure 40 is a sketch of a section of this coil showing the fin and
tube geometry. An arrangement with 14 fins to the inch, a face area 6 inches
wide by 12 inches high, and a 3-1/2 inch thickess was selected.
FIGURE 40. FIN AND TUBE GEOMETRY FOR
FUEL VAPORIZATION CHAMBER
-------
102
Figure 41 is a design layout of the vaporization chamber.
The
heating coil is shown mounted in a housing with inlet and outlet plenum chambers
to provide for uniform mixture distribution with minimum pressure loss.
The
coil is tilted at an angle from the axis of the inlet and outlet ports so as
to maximize the fin area exposed to the droplets. The housing would be
constructed with sufficient strength and rigidity to withstand maximum inlet
manifold vacuum. The inlet and outlet flanges of this configuration are
designed to be compatible with the proposed prototype induction system.
Simple
adaptors could be made to fit the vaporization chamber to conventional
induction system components.
The overall dimensions of this vaporization
chamber are:
height 16 in., width 7 in., and length 11 in.
Figure 42 is a sketch of the vaporization chamber mounted on a
351 CID V-8 engine with the prototype induction system components.
Shown
also in this sketch are the auxiliary components required.
In operation, hot
water is circulated through the coil by a pump.
The water is heated by the
addition of steam in a mixing chamber, the water absorbing the latent heat of
the condensing steam. Surplus water accumulating in the system by the
condensing steam will be drained off at the standpipe to maintain a constant
level in the circulation loop.
With an inlet temperature of 70 F, the system should be capable of
achieving a mixture temperature of 140 F at the outlet at full flow conditions.
This mixture temperature was selected for coil design from a consideration of
dew point temperatures expected in an operating engine. The dew point tempera-
ture is the lowest temperature at which all of the fuel will remain in vapor
form in equilibrium with the combustion air.
This temperature is a function
of the air-fuel ratio of the mixture and the manifold vacuum.
The mixture temperature at the inlet manifold will be controlled
by varying the temperature of the water being circulated through the coil.
A steam rate of about 36 lb/hr of saturated steam at atmospheric pressure
(about 4-1/2 gal/hr of water) will be required to vaporize the fuel and heat
the combustion air at full flow conditions.
A hot-water circulation rate of
4.0 gal/min will be used, giving a water-side pressure drop in the coil of
1.2 ft of water. The gas-side pressure differential across the coil at full
mixture flow of 300 cfm (at 2 in. Hg vacuum and 70 F) will be about 0.4 inches
of water.
-------
103
II"
..
Air and fuel vapor and droplets
rS.2"
16" !
I
12in.
\
\
\
\
\
\
\.
r
5.2"
Air and fuel vapor
, .
'\\\\
. \
\ \
\ \
\
\
~
~
ct
.1
~
\f1
I.
I'
/
\
liS-in. thick
steel plenum
""
\
\
....
....
(
V
I
V
f~
I
r
McQuay 5 BB 146 It 12.
hot water booster coil
(6-in. face width)
'\,
....
.~
/"
". 3-1/2 in. "-
/,,~'....
. \ " \
\ " \
\"""\
I' ~ \
\ ,,\.
"
, .
Coil support
pla1e
FIGURE 41. SKETCH LAYOUT OF VAPORIZATION
CHAMBER
-------
104
. !,...
Mixture generator
r
f fJ.
Overflow
10 droin
Air heater and vaporization
section (588 142-6X 12
McQuay heoting coil)
351 cu jn~ v-a engine
Standpipe
-e-- Steam
5 gollmin pump
FIGURE 42, FUEL VAPORIZATION APPARATUS
The system can be made to respond to a step increase in heating
load with little delay by greatly increasing the steam input. momentarily.
H~wever, the system is expected to respond to a step decrease in heating load
.' rather. slowly because heat is removed from the water only in the coil.'
Fift'e'eri m'inut~s or ,more may be required to achieve a stable inlet manifold
temperature at idle, following operation at full heating load.
ACKNOWLEDGEMENTS
This ~tudy was conducted under the sponsorship of the Office of
. 't... ,
Air Programs, Env~Jop~e~tal Protection Agency. The project Officer was Mr.
Jeffrey 1. Raney; o.r~gin~l1y and, later, Dr. Jose L. Bascunana of the
"'(, ) ;, .. ,
Characterizatio~ anCi'Co,nttol Development Branch of the Division of Emission
Contr61~Techno1ogy, with Mr. George D. Kittredge as Branch Chief.
-------
105
LIST OF REFERENCES
(1)
Trayser, D. A., et aI, '~ Study of the Influence of Fuel Atomization,
Vaporization, and Mixing Processes on Pollutant Emissions from Motor-
Vehicle, Power Plants", Phase Report from Battelle-Columbus Laboratories
to National Air Pollution Control Administration (April 30, 1969).
(2)
Fluid Meters - Their Theory and Application, Fifth Edition, The American
Society of Mechanical Engineers, New York City (1959).
(3)
Ranz, W. E., "Principles of Inertial Impaction", Department of Engineering
Research, Pennsylvania State University, Bulletin No. 66 (1956).
(4 )
Technical Data Book - Petroleum Refining, American Petroleum Institute,
Division of Refining, 1966 (with later revisions).
(5)
Obert, E. F., Internal Combustion Engines, Second Edition, International
Textbook Co., Scranton (1950), Chapter 8, pp 233-235.
(6)
Gieseke, J. A., and Mitchell, R. 1., "Size Measurement of Collected Drops",
J. Chern. Engr. Data, Vol 10 (No.4), p 350 (1965).
(7)
Doyle, A. W., Perron, R. R., and Shanley, E. S., "A Study of New Means
for Atomization of Distillate Fuel Oil", API Publication 1725 (1967).
(8)
Hazard, H. R., and Hunter, H. H., I~ Miniature Ultrasonic Burner for a
Multifueled Thermoelectric Generator", API Conference Paper CP66-3,
API Publication 1705 (1966).
(9)
Martner, J. G., "An Ultrasonic Atomizer Capable of High Rates", API
Conference Paper CP66-5, API Publication 1705 (1966).
(10) Hartmann, J., and Trolle, B., "A New Acoustic Generator, The Air-Jet
Generator", Journal of Scientific Instruments, Vol 4 (No. 101), 1926-27.
(11) "Development of a Low Emission Burner for Rankine-Cycle Engine", EPA
Contract No. EHS 70-117.
(12) Personal correspondence with Sonic Development Corporation of America.
(13) Mitchell, R. 1., and Pilcher, J. M., "Improved Cascade Impactors for
Measuring Aerosol Particle Sizes in Air Pollutants, Commercial Aerosols
and Cigarette Smoke", Ind. & Eng. Chemistry, Vol 51, No.9 (1959).
-------
APPENDIX A
IMPACTOR STAGE DESIGN FOR VAPOR/DROPLET SAMPLERS
-------
APPENDIX A
IMPACTOR STAGE DESIGN FOR VAPOR/DROPLET SAMPLERS
The design procedures for the three impactor geometries (vapor/droplet
samplers) were all similar. The design was based on known and measured efficiencies
for a similar impactor design. The relationship governing the design of an
impactor stage can be expressed as
dj ~
p V.D 2 = KI
P J P
(A-I)
where
dj = diameter of jet
Pp = density of droplet liquid
V. = gas velocity through the jet
J
D = droplet diameter
p
Kr = impactor design constant
~ = gas viscosity
A further consideration is that the distance from the jet to the impaction
surface should be about 0.4 times the" jet diameter.
The value of the constant, KI' in Equation A-I is available from the
design for the Battelle cascade impactor (12) which has been calibrated. This
value can be taken as
-11
KI = 6.92 x 10 in/lb
The density of Stoddard
Solvent is 49.3 lb/ft3 and the maximum velocity in the
test sections was expected to correspond to 100 scfm through a cross section
of 1.25 x 1.25 inches giving a velocity of 164 ft/sec at atmospheric pressure.
The sampling rate was chosen to be 0.71 scfm which gives a velo~ity of 167 ft/
sec through a jet with a diameter of 0.113 inch. If this jet diameter is
chosen, the sampling probe under the highest expected flow conditions would
-------
A-2
be the same size as the jet.
This jet size was therefore chosen for these
reasons and also because the cut-off size for the impactor would be 1.35 microns
as calculated with Equation A-I which was assumed to be a small enough droplet
size for the purposes of separating drops from the vapor.
The vapor/droplet samplers were calibrated by measuring the size
distribution of drops passing through the impactor on the gas side. The
calibrations were performed using water droplets containing a dissolved
fluorescent tracer (sodium fluorescein dye). The water was atomized from a
DeVilbiss D-40 atomizer operated with an air supply at 5 psig. The DeVilbiss
D-40 atomizer is a simple dip-tube aspirating atomizer available at many drug
stores for use in inhalation therapy. The spray was directed into a duct four
inches in diameter and two feet long, and two-minute samples were taken from
the duct exit with the vapor/droplet sampler. The size distribution of droplets
produced by the atomizer is presented as Figure A-I. The data points on this
graph are averages of five measurements made with the Battelle cascade impactor.
Measurements of sampling flow rates were made with calibrated
rotameters. The total sampling rate was 0.75 scfm, with each half, vapor and
droplet, being 0.375 scfm. The gas flow through the vapor side was sampled
into a cascade impactor to determine the droplet size cutoff point. On the
droplet size both the liquid collected in the sampler and that collected on a
backup filter following the sampler were measured. All measurements were of
fluorescence, which is proportional to liquid mass, and were made with a
fluorophotometer.
The mathematical procedure for determining the impactor efficiency
can be explained in terms of mass flows of droplets of different sizes.
Figure A-2 illustrates the sampler calibration arrangement and the mass flows
of droplets. The desired information to be derived is the efficiency of the
vapor/droplet sampler, E., for different droplet sizes noted by the subscript i.
~
This f~actional efficiency can be defined as:
E. = m./M.
~ ~ ~
where
Mi = mas's of droplets of size i entering the vapor/droplet sampler
mi = mass of droplets of size i collected in the vapor/droplet sampler.
-------
A-3
99
98
95
90
80
- 70
.r:.
01 60
cu
~
~
- 40
c
cu 30
u
L..
tf
cu 20
>
-
0
::J
E
::J
U 5
/
- /
v
/
vi
/
V
V
, /
/
V
1--- /
/
)
/
) v
v
V
./ V
/'
50
10
2
0.5
0.2
0.1
0.05
0.0'0.2
0.3 0.4
0.6 0.8 1 2 3 4
Equivalent Particle Diameter, microns
6
8
10
20
FIGURE A-I. DROPLET SIZE DISTRIBUTION USED FOR CALIBRATION
OF THE VAPOR/DROPLET SAMPLER
-------
A-4
Mj
Toto I --.
sample
Vapor/droplet
sampler
Vapor" flow, Mj - mj
- Cascade impactor
"l
Droplet
flow, Lmj
FIGURE A-2. VAPOR/DROPLET SAMPLER CALIBRATION ARRANGEMENT
The cascade impactor separates the droplets into size fractio~s (with average
size i) with the mass fraction of each size noted by xi = Mi/EMio When the
cascade impactor is preceded by the vapor/droplet sampler the measured
fraction in each size range is given by Yi = (Mi - mi)/E(Mi - mi)o
this equation gives:
Rearranging
mi = 1 - Yi E(Mi - mi)
Mi Mi""
(A-3)
or
Yi E(Mi - mi)
e; = 1 -
i xi EMi
(A -4 )
The total mass of all sizes sampled, EM., is measured as the total mass
1
collected in the vapor/droplet sampler and the cascade impactor, or
EMi :;: EMi + I(Mi - mi)
(A-S)
Then in values measured experimentally,
-------
A-5
Y i E (Mi - m.)
e: = 1 - (-) [ 1. ]
i xi Emi + E(Mi - mi)
(A-6)
The experimental calibration procedure is to sample with the cascade
impactor alone to obtain xi' Then the vapor/droplet sampler is inserted ahead
of the cascade impactor and the droplet distribution sampled again. From this
sample values for Yi' E(Mi - mi)' and Emi are determined and the efficiency,
e:i' calculated.
F~gure A-3 shows the efficiency of the vapor/droplet sampler as a
function of droplet size for sample'!: Designs 2 and 3. .The fact that the
fractional efficiency does not go below 0.5 reflects the sampling situation
where one-half of the total vapor and small-droplet flow passes through the
impaction stage along with any larger droplets. In other words, even if there
is no inertial impaction of the smaller droplets they are carried along with
the larger droplet flow. Impaction is seen to be 90 percent effective for
drops of 2 microns, and more efficient for larger drops.
efficient than predicted, it is still sufficient.
Although this is less
-------
A-6
(1,-, : -
1.0
~ H),
7 V~
1/ .A1~
V
~ ('
/
V "
--'
0.8
>-
u
c:
Q,)
.~ 0.6
-
-
I.LI
o
c:
.~ 0.4
-
u
o
~
I.L..
0.2
0.00
.1
0.2 0.3 0.4 0.6 08 I 2
Drop Dio meter, microns
3 4
6
8 10
FIGURE A-3. COLLECTION EFFICIENCY OF THE VAPOR/DROPLET
SAMPLER AS A FUNCTION OF DROPLET SIZE
-------
APPENDIX B
CALIBRATION CURVES
-------
APPENDIX B
CALIBRATION CURVES
The significance of the curve sheets presented in this Appendix
should be self explanatory, with the exception of Figures B-3 through B-14.
These figures indicate the droplet sizes produced by the spinning-disk.
atomizer, used in the laboratory apparatus, in free air, as a function of
fuel feed rate and disk speed. Actual drop-size distribution entering the
test section was somewhat smaller, as shown in Figure 22.
I -
-------
B-2
200
2
"
/" /'
/
~
./
.~
, ,g;,
e~/ ~
~~ /'"
. ce
O(\,~ /"
./ /"
./ / 'Ooo'~ ~
~ o.
./ '
e~
.;' 6\o~
O(\,\c~~
[//
V ~
/ ./
/ /'"
./
\~ I
. 0& I
,.
e'~
6~/
've./
O~r
/", '"
/'
/ V
---:;7
100
80
60
40
30
c:
N
~ 20
en
N
IJ...'"
en
I{)
-
o
E
-
u
en
10
8
~
o
IL.
~
6
~
4
3
I
I
2
3 4 6 8 10 20
Pressure Differential, in. H20 at 75 F
30
40
60
FIGURE 8-1. COMPUTED FLOW VERSUS PRESSURE DIFFERENTIAL DATA FOR
SQUARE-EDGED ORIFICE FLOW METER WITH I-D AND 1/2-0 TAPS
-------
'B-3
200
25
,) I
/
J
Stoddard solvent \7
7
/
/r
/
/
) /
175
150
125
c
-€
u
u
i 100
o
lL.
(1)
:3
u...
75
50
00
2
4
6 8 10
Rotameter Reading
12
14
16
18
FIGURE B-2. CALIBRATION CURVE FOR FUEL SYSTEM
FLOW METER
'.
-------
B-4
99
-
.Q
='
E
='
u
5
b
j
r
- 0/
7
19
I
- f
i
/
I
/
r
98
95
90
80
70
~ 60
0'
'0) 50
;=
>- 40
.0
C 30
C1>
o
l-
ff 20
C1>
:>
10
2
0.5
0.2
0.1
0.05
0.012
3
4
6 8 10 20 30 40
Equivalent Particle Diameter, microns
60
80 100
FIGURE B-3. DROP SIZE DISTRIBUTION AT 28.000 RPM AND
I CM3/SEC
-------
99
98
95
90
80
70
- 60
.c
.Q'
Q) 50
~
>- 40
.Q
-
c 30
Q)
u
'-
Q) 20
a..
Q)
>
- 10
o
::J
E
::J 5
u
2
0.2
0.1
0.05
B-S
i
---- -..
A
11
r
cj
-- /
/
---
-.. p
j
/
!
/
~
y ./
0.5
0.012
3
4
6 8 10 20 30 40
Equivalent Particle Diameter, microns
60
80 100
FIGURE B-4. DROP SIZE DISTRIBUTION AT 40,000 RPM AND
I CM3/SEC
-------
B-6
99
2
rI
l6
6
r!
/
j
I
j .
, /
I
j
/
J
,)
98
95
90
80
70
-
&. 60
.2'
~ 50
~ 40
-
~ 30
e
~ 20
CI)
>
- 10
o
-
::J
E
::J 5
u
0.5
0.2
0.1
0.05
0.012
3
4
6 8 10 20 30 40
Equivalent Particle Diameter, microns
60
80 100
FIGURE B-5. DROP SIZE DISTRIBUTION AT 46,000 RPM AND
I CM3/SEC
-------
B-7
99
98
2
/...
,...,
?
/
9
/
V
,.J
J
j
/ .
0/
/
/
J
V
I ~
1----
95
90
80
70
- 60
~
01
'i) 50
~
>- 40
.0
'E 30
Q)
~
~ 20
Q)
.~
o 10
"5
E
::J 5
u
0.5
0.2
0.1
0.05
0.01
2
3
4
6 8 10 20 30 40
Equivalent Particle Diameter, microns
60 .80 100
FIGURE 8-6. DROP SIZE DISTRIBUTION AT 53,000 RPM AND
I CM3/SEC
-------
B-8
70
6
,I
)
/
') /, 0
-/
J
cI
/
7
h( l/
1/
3
4
6 8 10 20 30 40
Equivalent Particle Diameter, microns
60 80 IPO
99
98
95
90
80
:E 60
CI
'Q) 50
~
>- 40
.0
- 30
c
fj
~ 20
a..
CD
>
- 10
o
::J
E 5
::J
U
2
0.5
0.2
0.1
0.05
"0.012
FIGURE B-7. DROP SIZE DISTRIBUTION AT 59,000 RPM AN D
I CM3/SEC
-------
B-9
99
98
95
90
80
70
- 60
~
01
CI)
~
>-
.0
-
C
CI)
u 20
L.
If
CI)
>
-
0
::I
E
::I
U
I
I
/
/ J
/
/
/
f
V
v
Ij
I
50
40
30
10
5
2
0.5
0.2
0.1
0.05
0.012
3
4
6 8 10 20 30 40
Equivalent Particle Diameter, microns
60
80 100
FIGURE B-8. DROP SIZE DISTRIBUTION AT 64,000 RPM AND
I CM3/SEC
-------
J
?
/
9
~
I
P
(f
I
f
J
1
1/
/
J
I
I ,)
99
98
95
90
80
70
60
.....
~
0' 50
'Q;
~ 40
>-
..0
..... 30
c:
cu
e 20
cf
cu
.~ 10
-
c
='
E 5
='
u
2
0.2
0.1
0.05
0.012
0.5
3
4
B-10
6 8 10 2D 30 40
Equivalent Particle Diameter. microns
60 80 100
FIGURE B-9. DROP SIZE DISTRIBUTION AT 32.000 RPM AND
0.5 CM3/SEC
-------
B-ll
99
98
,
r
j
/
))
j5
0/
/
I
I
I
1/
I )
1/
J
/
V
2 3 4' 6 8 10 20 30 40 60 80 1
95
90
80
70
60
50
-
s:;
C'
cu
~
-
c:
cu
u
'-
cf
cu
>
-
0
:::J
E 5
:::J
U.
>- 40
.0
30
20
10
2
0.5
0.2
0.1
0.05
0.01
00
Equivalent Particle Diameter, microns
FIGURE B-IO. DROP SIZE DISTRIBUTION AT 43,000 RPM AND
0.5 CM3/SEC
-------
B-12
99
~8
95
I
/
o /
-
/
!
/
5
/
j
r
J
IV
I
/
/
P
1/
3
4
6 8 10 20 30 40
Equivalent Particle Diameter, microns
60
80 100
90
80
70
- 60
~
.~ 50
cu
~
>- 4.0
.0
-
c
cu
u
'-
cu
Cl.
cu
>
- 10
o
:::s
E 5
:::s
U
2
30
20
0.5
0.2
0.1
0.05
0.012
FIGURE B-II. DROP SIZE DISTRI BUTION AT 50,000 RPM AND
0.5 CM3;SEC
-------
B-13
99
/ 0
J
7
/
'd
/
/
f
/
jd
V
/
"1
J
98
95
90
80
70
.... 60
.s::.
.2' 50
Q)
~
40
>-
.c
....
c
Q)
u
'-
cf
(I)
.2:
....
o
"3
E
::J
U
30
20
10
5
2
0.5
0.2
0.1
0.05
0.012
6 8 10 20 30 40
Equivalent Particle Diameter, microns
60 80 100
3
4
FIGURE B-12. DROP SIZE DISTRIBUTION AT 56,000 RPM AND.
0.5 CM3/SEC
-------
"
B-14
99
-
.s::;
C'
V
~
.
1>
/
8
/
J
r
/
/
t .
/
I
(J
I
I
J
V
/
(l{
/
98
95
90
80
70
60
>-
.c
50
40
-
c:
v
U
l-
V
ll.
v
>
30
20
10
-
o
::s
E
::s
u
5
2
0.5
0.2
0.1
0.05
0.012
4
6 8 10 20 30 40
Equivalent Particle Diameter
60 80 100
3
,FIGURE B-13. DROP SIZE DISTRIBUTION AT 61,000 RPM AND
0.5 CM3/SEC
-------
1_-
B-15
99
98
-
~
01
~
I
9
/
/
/
/
?
/
/ /
1/
V
oJ
I J
95
90
80
70
60
50
>-
.0
40
30
-
c:
Q)
u
....
Q)
Cl.
Q)
>
20
-
o
::J
E
::I
U
10
5
2
0.5
0.2
0.1
0.05
0.01
2
3
6 8 10 20 30 40
EQui va lent Particle Diamete r. microns
4
60
80 100
FIGURE B-14. DROP SIZE DISTRIBUTION AT 65,000 RPM AND
0.5 CM3/SEC
-------
150
125
100
en
~
Q)
...
Q)
E
E 75
01
c:
~
0 t:I:'
Q) I
a: .....
50 0-
~
Q)
...
Q)
E
0
...
0
a:
25
00
5
10 15 20 25 30
Air Flow~ standard liters per minute at 75 F. 14.7 psia
40
35
FIGU !E B-15. STANDARD CALIBRATION OF PRIMAR' SAMPLING ROTAMETERS
-------
1.0
\
'\ 6. '8' rotameter
0, 00 'A' rotameter
'
" ~
~ "--
u............... ~
- "
L1'--- ----
'-'
C::. IV"
~
I
......
"'-J
0.8
'-
o
-
g 0.6
lL.
c
o
~
u
CD
t 0.4
o
U
0.2
0.0 0
I.
2.0
3.0
4.0 5.0 6.0 7.0. 8.0
Correction Para meter ( 460 + T~ of )
t 530
9.0
(. 29.92 )
,P. in. Hg
10.0
11.0
12.0
FIGURE 8-16.
TEMPERATURE AND PRESSURE CORRECTIONS FOR SAMPLING ROTA METERS
-------
35
~ 30
CI)
-
::::
en
~
CI)
~ 25
0
~
u
E
en 20
0
C>
c
c 15
0
-
0
~
-
c
CI) tJ:j
u 10 I
C ......
o 00
u
CI)
::I
I.L. 5
20
40
60
80 100 120 140 160
Hydrocarbon Analyzer Reading x 10-3
180
200
220
FIGURE B-17.
HYDROCARBON ANALYZER CALIBRATION
------- |