DOE/NBB-0067

DOE/NASA/0143-1

NASA CR-174813

Research and Development of Neat

Alcohol Fuel Usage in Automobiles

June 1984

,

(NASA-CB-174813) BESIABCH ANC DEVELOPMENT

OP NEA T ALCOHOL FDEL. DSAGE IN AD10MOBILES

(Santa Clara Univ.^. 147 p HC A07/MF A01

CSCL 21D

I G3/28

>,

N86-27460

Unclas

43119 J

Prepared for

National Aeronautics and Space Administration

Lewis Research Center

Under Grant No. NAG 3-143

for

U.S. Department of Energy

Assistant Secretary for Conservation

and Renewable Energy

Office of Vehicle and Engine R&D

Under Interagency Agreement DE-A101-77CS51044

DOE

 Alternative

Alcohol Fueis

Fuels

 Utilization

pr°9ram Program
DISCLAIMER

"This report was prepared as an account of work sponsored by an agency of the United

States Government. Neither the United States.Government nor any agency thereof, nor any

of their employees, makes any warranty, express or implied, or assumes any legal liability or

responsibility for the accuracy, completeness, or usefulness of any information, apparatus,

product, or process disclosed, or represents that its use would not infringe privately owned

rights. Reference herein to any specific commercial product, process, or service by trade

name, trademark, manufacturer, or otherwise, does not necessarily constitute or imply its

endorsement, recommendation, or favoring by the United States Government or any agency

thereof. The views and opinions of authors expressed herein do not necessarily state or

reflect those of the United States Government or any agency thereof."

This report has been reproduced directly from the best available copy.

Available from the National Technical Information Service, U. S. Department of Commerce, Springfield,

Virginia 22161.

Price: Printed Copy AQ8

Microfiche A01

Codes are used for pricing all publications. The code is determined by the number of pages in the

publication. Information pertaining to the pricing codes can be found in the current issues of the following

publications, which are generally available in most libraries: Energy Research Abstracts, (ERA);

Government Reports Announcements and Index (GRA and I); Scientific and Technical Abstract Reports

(STAR); and publication, NTIS-PR-360 available from (NTIS) at the above address.
DOE/NBB-0067

DOE/NASA/0143-1

NASACR-174813

Distribution Category UC-926, UC-96

Research and Development of Neat

Alcohol Fuel Usage in Automobiles

June 1984

Prepared bv:

R.K. Pefley, Project Director

Co-Authors: L. Browning, S. Espinola,

B. Pullman, R. Gururangan, N. Saito

Mechanical Engineering Department

University of Santa Clara

Santa Clara California 95053

Under Grant No. NAG 3-143

Prepared for

National Aeronautics and Space Administration

Lewis Research Center

Cleveland, Ohio 44135

for

U.S. Department of Energy

Assistant Secretary for Conservation

and Renewable Energy

Office of Vehicle and Engine R&D

Washington, D.C. 20585

DOE Alternative

Alcohol Fuels

Fuels Utilization

Program Program
INDEX

PAGE

List of Figures i

List of Tables iii

I. INTRODUCTION I.I

II. COMPUTER MODELING OF- ENGINE PROCESSES II.1

1. COMPUTER AIDED DESIGN STUDY FOR POST OFFICE ALCOHOL II.1

VEHICLES

a. Fuel Type Comparisons I I.I

b. Compression Ratio Comparisons II.2

c. Rear Axle Ratio Comparisons II.2

d. Reduction of Parasitic Cooling Loads II.3

e. Summary 11.3

2. NOx CONTROL TECHNIQUES FOR METHANOL FUELED SI ENGINES II.4

a. The computer Model II.4

b. Spark Retard Effects 11.6

c. Compression Ratio Effects II.6

d. EGR Effects 11.11

e. Summary 11.11

3. COLD START ENGINE PROCESS MODELING 11.14

a. Droplet Fall-Out and Sling-Out Model 11.17

b. Compression Process Model 11.20

c. One-Dimensional, Two-Phase Cold Start Engine Model 11.29

d. Summary 11.35

4. REFERENCES 11.36

III. ENGINE COLD STARTING BY EXPERIMENTATION 11 I.I

1. COLD START PROBLEM ELEMENTS 11 I.I

2. EXPERIMENTAL APPROACH III.2

3. THE EXPERIMENTAL HARDWARE III.3

4. THE DATA ACQUISITION SOFTWARE III.7

5. THE TEST MATRIX AND OTHER EXPERIMENTAL RESULTS III.9

a. Experimental Determination of Blowby Area III.10

b. Determination of Valve Dishcarge Coefficient III.14

6. DATA REDUCTION AND ANALYSIS 111.14

7. MODEL VALIDATION III.17

8. SUMMARY III.21

9. REFERENCES III.22
"". r - PAGE

IV. VEHICLE AND ALLIED ENGINE STUDIES . . IV.1 ' , • t • • • „ *•

, 1. POST OFFICE PROJECT INTRODUCTION IV.1

a. Vehicle Modification Criteria IV.1

1. Power and Fuel Economy " IV.2

2. Materials Compatibility ...;, IV.2

b. Cold Startability and Driveability IV.3

•_;- c. Installation of Kits and Field Experience . IV.4

1. Field Experience IV.4

d. Summary IV.5

- 2. EGR STUDY IV.6

.- = a. EGR Limits IV.6

b. Variable Load Tests at Constant Speed IV.6

c. Variable Speed Tests at Constant Load IV.7

d. Conclusions j IV.7

3. UNBURNED FUEL EMISSIONS CHARACTERIZATIONS IV.8

a. Description and Results of the Blend Tests IV.13

b. Description and Results of the Neat Fuel Studies IV.14

c. Concluding Observations IV.16

4. TECHNIQUES FOR EVALUATION OF EMISSIONS OF TESTING IV.16

ALCOHOL FUELED VEHICLES

a. Unburned Fuel Emissions IV.16

b. Fuel Economy Calculations IV.17

c. Water Condensation IV.18

d. Conclusion IV.18

5. TOYOTA CRESSIDA—A UNIQUE TEST VEHICLE IV.18

a. Introduction IV.18

b. The Toyota Closed Loop Feedback Controlled IV.19

Fuel Injection System

c. Test Results and Discussion —1980 Cressida IV.21

d. Rise of NOx Emissions From Alcohol Blends IV.22

e. Operating Experience IV.23

f. Test Results and Discussion — 1982 Cressida IV.24

g. Operating Experience IV.25

h. Conclusions IV.30

6. REFERENCES IV.30
PAGE

V. CORROSION AND WEAR MECHANISMS

 : V.I

1. A PROPOSED WEAR MECHANISM ' V.3

2. MEASUREMENT OF THE CORROSION POTENTIAL AND CORROSION V.5

RATE

3. MEASUREMENT OF THE CONDENSATE DIFFUSION RATE V.ll

4. FURTHER ANALYSIS OF THE CONDENSATE V.I3

5. CONCLUSIONS AND FUTURE WORK ' V.16

6. REFERENCES . V.18

VI. PHOTOCHEMICAL SMOG CHAMBER STUDIES VI.1

1. FACILITY UPGRADE ' ' VI.1

. 2. INSTRUMENTATION VI.2

3. CHAMBER GAS SAMPLING VI.3

4. CALIBRATION STUDY VI.4

5. EXPERIMENTATION VI.17

6. CONCLUSIONS VI.19

' 7. REFERENCES VI.23
List of Figures

Chapter Page

11.1 Spark Retard Effects II.7

11.2 Compression Ratio Effects Constant Open-to-Squish II.9

Volume Ratio

11.3 Compression Ratio Effects Constant Squish Clearance 11.10

Height

11.4 EGR Effects 11.13

11.5 Saturated Vapor/Air Equivalence Ratio vs. 11.15

Temperature for Methanol

11.6 Saturated Vapor/Air Equivalence Ratio vs. 11.16

Temperature for Ethanol

11.7 Droplet Fall-Out and Sling-Out Model Schematic 11.18

11.8 Droplet Size Effects on Droplet Trajectory 11.19

11.9 Cranking Speed Effects on Droplet Trajectory 11.21

11.10 Cylinder Pressure for Room Temperature Cranking 11.22

11.11 Cylinder Temperature for Room Temperature Cranking 11.24

11.12 Fuel Effects for Room Temperature Cranking 11.25

11.13 Fuel Effects on Minimum Starting Temperature 11.26

11.14 Droplet Size Effects on Minimum Starting Temperature 11.27

11.15 Inducted Equivalence Ratio Effects on Minimum Starting 11.28

Temperature

11.16 Cranking Speed Effects on Minimum Starting 11.30

Temperature

11.17 Compression Ratio Effects on Minimum Starting 11.31

Temperature

11.18 Manifold Vacuum Effects on Minimum Starting 11.32

Temperature

11.19 Starting Temperature Effects 11.34

111.1 Cold Start Experimental Set-Up III.4

111.2 Head Instrumentation III.5

111.3 Shaft Encoder Assembly III.5

111.4 Schematic of Experimental Set-Up III.6

111.5 Flow Chart for the Data Acquisition Software III.8

111.6 Pressure Crank Angle History III.11

111.7 Pressure Drop at Different Temperatures III.12

1II.8 Cylinder Blow-By 111.13

111.9 Intake Valve Lift During Cranking III.15

111.10 Intake Valve Discharge Coefficient III.16

111.11 Heat Flux, Model vs. Experiment III.18

111.12 Cylinder Temperature, Model vs. Experiment III.19

111.13 Cylinder Pressure, Model vs. Experiment III.20

IV.1 Unburned Fuel Emissions from Neat or Near Neat IV.11

Alcohol Fueled Vehicles

IV.2 Unburned Fuel Emissions from Alcohol/Gasoline IV.12

Blend Fueled Vehicles
Chapter Page

IV.3 Toyota EFI Closed Loop Control System IV.20

IV.4 FTP CO Emissions vs. Methanol Blend IV.26

IV.5 FTP HC Emissions vs. Methanol Blend IV.27

IV.6 FTP NOx Emissions vs. Methanol Blend IV.28

IV.7 FTP NOx Emissions vs. Methanol Blend with IV.29

Spark Retard

V.I Electrochemical Cell V.8

V.2 Corrosion Potential vs. pH V.9

V.3 Typical Anodic Polarization Curve V.10

V.4 Correlation of Oil Coating Film Thickness with V.12

Rotation Speed

V.5 Transition Time vs. Oil Film Thickness for V.14

Condensate with Various Methanol Percentages .

V.6 Exhaust Condensate pH Values for Various V.15

Vehicle/Fuel Configurations

VI.1 Experimental vs. Calculated SMOG Chamber Dilution VI.6

Comparison

VI.2 Typical Temperature Profile During SMOG Chamber VI.7

Experiment

VI.3 Acetaldehyde Characterization to Measure NOx VI.9

Offgassing Rate

VI.4 Characterization Experiments Using Clean Air VI.10

and NOx

VI.5 Characterization Experiments Using Carbon VI.11

Monoxide and NOx

VI.6 Propene & NOx Characterization for Formaldehyde VI.13

Photolysis

VI.7 Experimental & Urban Formaldehyde Dioxide Photolysis VI.14

Rate Constants

VI.8 Experimental & Urban Nitrogen Dioxide Photolysis VI.16

Rate Constants

VI.9 Reactant Concentrations and SMOG Chamber Products VI.18

Over Time

VI.10 Baseline Surrogates, 2/3 Baseline & Methanol VI.20

Substitution at 3:1

VI.11 Baseline Surrogates, 2/3 Baseline & Methanol VI.21

Substitution at 9:1 *

VI.12 Baseline Surrogates, 2/3 Baseline & Methanol VI.22

Substitution at 27:1

ii
List of Tables

Chapter Page

III.l Test Matrix III.9

IV.1 Detailed UBF Exhaust Emissions Summary IV.9

IV.2 Detailed UBF Evaporative Emissions Summary IV.10

V.I pH and Acid-Base Titrations of Pre-Catalyst V.2

Exhaust Condensate

V.2 Composition of the Electrolytic Solutions V.6

V.3 Ion Concentrations (mg/L) and pH Values of V.16

Pre-Catalyst Exhaust Condensate

VI.1 Urban Hydrocarbon & Methanol Exhaust Surrogates VI.15

VI.2 Experimental Design & Initial Surrogate VI.17

Concentrations

ill
CHAPTER I

INTRODUCTION

This grant has provided a unique opportunity to start the bridge between

researching the basic feasibilities of alcohol (methyl, known hereafter as

methanol and ethyl, known as ethanol ) and commercializing these fuels as

extenders and supplanters of petroleum in the transportation industry. A bridge

is suggested because the research at Santa Clara started in 1968 with a National

Air Pollution Control Agency (NAPCA) project investigating the clean burning

aspects of methanol to reduce photochemical smog and ranged in 1979 to DOE

sponsored alcohol fueled vehicles being demonstrated at the DOE Contractors

Coordination Meeting in Detroit; on Capital Hill and at the DOE, Forrestal

Building, Washington, D.C. in November of 1979. During that time frame four

international alcohol fuel symposia as well as domestic meetings clearly

established methanol and ethanol as the most likely near term alternatives to

petroleum. They were shown to allo w common power plants to continue to be

produced throughout the transportation world, (the spark ignited, SI, and

compression ignited, CI, engines) yet, al low a large number of countries to

preserve or restore their liquid energy independence because they can be

produced for coal, natural gas, tar sands, residual crude, shale oil and

biomass.

This grant ranges in objectives from nearly basic research in the form of

computer modeling of reactive thermofluid mechanics to commercial aspects where

after-market conversion kits were engineered to change Post Office vehicles from

gasoline to methanol and ethanol fuels. During this time frame pure alcohol

fleet vehicle numbers have grown in the U.S from the range of 1-10 through the

range of 10-100 and now are in the range of 100-1000.

While alcohol fuels have moved towards commercialization, two problems have not

been satisfactorily resolved. These are: cold-start, warm-up driveability and

corrosion and wear of fuel system and engine parts. Additionally, environmental

factor assessments must continuously be made if the public is to be assured that

the alcohols are more beneign than petroleum in terms of environmental, health,

and safety factors.

I.I
This report is a synopsis of the work; designed to push the alcohols along the

path of acceptance, for it is bel ieved that properly engineered they will allow

the world finally to balance its transportation energy requirements with the

energy harvested from the sun, thereby freeing transportation energy from

petroli'eum^and oppressive prices associated with monopolies.

 :

The report is heavily keyed to the references which include all of the

publications associated with the grant. In effect, it is a summary of-this

evidence which may • be found in greater detail elsewhere in the literature; ;:

1.2
--•'. ;,<-'• - CHAPTER II c< ' . ..

COMPUTER NODELING OF EHGINE PROCESSES

Computer modeling of complex engine processes has been successfully used in

research investigations at the Uni versity of Santa Cl ara to provide a ,better

understanding of trends in engine performance and pollutant emissions produced

from our laboratory tests and found in the literature. By using a mix between

theory and.empiricism, these models have greatly added to our knowledge about

various aspects of.utilizing alcohols in automotive engines. Unless noted

otherwise, all modeling is based on the physical dimensions of the 2.3 1 iter

engine as found in Ford Pinto vehicles.

II.1 COMPUTER AIDED DESIGN STUDY FOR POST OFFICE PROTOTYPE ALCOHOL

VEHICLES

A computer model that simulates the second-by-second driving cycles, such as the

Federal Test Procedure, was used in this study to examine the effects of changes

in fuel type, compression ratio, rear axle ratio and reduction of parasitic

loads on energy economy and exhaust emissions. Three driving cycles were

defined for this study: the urban driving cycle, the highway driving cycle and

the Post Office delivery cycle. Steady-state performance and emissions maps

taken from our engine-dynamometer facility were used as inputs to this model.

The model, however, does not consider engine warm-up periods or transient

performance. The model has been shown to yield economy figures comparable to

actual vehicle tests; however, the emissions figures are most useful when

compared to other simulated runs [II.I].*

II.la Fuel Type Comparisons

Three fuels, methanol , ethanol and gasoline, were compared for all three driving

cycles. Methanol showed substantial improvements in energy economy over

gasoline in all tests. At an average equivalence ratio** of 0.9, methanol

showed 18%, 14% and 26% improvement over gasoline for the urban, highway and

* [ ] Denotes references at the end of the Chapter.

** Stoichiometric air/fuel (A/F) ratio divided by actual air/fuel ratio.

II.1
Post Office driving cycles respectively. Ethanol showed 11%, 12% and 11%

improvement over gasoline for the same comparisons. The engine exhaust NOX

emissions .from methanol were considerably lower than for gasoline1under the same

conditions, while the other emissions remained about the same. Ethanol, however,

showed-a substantial increase in hydrocarbon emissions and a substantial

decrease in CO emissions while NOY emissions remained about the same as for

/\

gasoline.

Il.lb Compression Ratio Comparisons

Increasing the compression ratio seemed to have little effect on energy economy

for methanol. Raising the compression ratio from 8.5 to 11.3 increased the

energy economy 9%, 14% and 5% for the urban, highway and Post Office Driving

cycles. Since raising the compression ratio increases power, a lower throttle

setting must be used to maintain the same power level at a given speed. By

doing this, the engine is forced to operate at a lower thermal efficiency due to

increased exhaust dilution which can offset thermal efficiency gains from

raising the compression ratio. Due to the faster burn rate of the 11.-3 compression ratio engine, NO emissions were shown to be 31% lower than the 8.5:1 A

engine [2], due to MBT timing closer to TDC and less time at high temperatures

and pressures for NOX formation. Hydrocarbon emissions however rose 118% over

the standard 8.5:1 engine. •

II.lc Rear Axle Ratio Comparisons "

As stated in the previous section, raising compression ratio without increasing

the load on the engine results in decreased efficiency. By lowering the rear

axle ratio while raising compression ratio, the load on the engine is increased

for the same power output at the wheels. Using methanol in the 11.3 compression

ratio engine, it was found that lowering the rear axl e ratio from 3.08 to 2.73

gave; an additional fuel savings of 5%, 12% and 7% respectively for the urban,

highway and Post Office driving cycles.

II.2
11. Id Reduction of Parasitic Cooling Loads

Since the alcohols have much higher heats of vaporization than gasoline,- the

internal cooling effects of fuel vaporization in the cylinders is. much greater.

With this in mind, an alcohol engine can be designed with a smaller, cool ing

system thereby reducing the parasitic cooling loads. The easiest way of accomplishing this would be to disconnect the cooling fan and water pump from the

engine and drive them electrically, on an as-needed basis. In order to see the

effects of this modification, the model was run with and without the cooling fan

and water pump in the accessory loads. Again using the 11.3 compression ratio

engine with methanol, 6%, 5% and 3% gains in energy economy can be realized on

the urban, highway and Post Office driving cycles respectively by removal of the

fan only, while an,-8%, 7% and 5% gain can be realized for removal of the fan and

water pump.

Il.le Summary

The computer-aided design study presented here indicates'gains in fuel economy

and reductions in emissions by all of the alcohol conversions considered. The

computer model suggests that conversion to methanol yields the best energy

economy. Raising compression ratio shows substantial fuel economy benefits but

results in a HC emission penalty. The rear axle ratio change and removal of the

engine cooling components shows slight fuel economy and emissions benefits but

should be weighed against conversion costs.

In the future, alcohol fueled engines will be smaller than their gasoline

counterparts enabling them to more effectively exploit the properties of the

alcohols and derive the anticipated fuel economy.

As a last word of caution, it should be noted that the data taken for the

•driving cycle simulations was taken from a fully warm, dynamometer-mounted

engine operated at steady-state conditions and did not include transient engine

behavior or warm-up performance and emissions. These factors need to be taken

into account before a final decision is made on the modification package for a

vehicle with a gasoline engine.

II.3
II.2 NOX CONTROL TECHNIQUES FOR METHANOL FUELED SI ENGINES

While alcohol fuels are generally clean burning fuels, air pollution problems

existing in major cities throughout the world will eventually require emissions

controls on all combustion sources. Since oxides of nitrogen (NOX) emissions

are fundamental in the formation of photochemical smog and the hardest emissions

to control, a review of NOY control techniques is presented. With this- study, A

NOY control methods can be considered without large penalties in performance or A

other exhaust emissions and experimental costs.

The method used to study these NOV control techniques is a fundamental spark A

ignition (SI) engine computer model. This approach is used since the interaction between engine design and operating variables is sufficiently complex

that a fundamental model is essential in interpreting performance and emissions

results from a spark ignition engine. A fundamental model employs current

knowledge and semi-empirical formulations of combustion and pollutant formation

processes occurring during an SI engine cycle to predict, in a reasonably

expedient and low-cost fashion, the results of various changes in engine design

and operating variables. The results from this kind of study can be used to

guide experimentalists in interpreting their results and engine designers in

screening various concepts in engine design prior to expensive hardware development and testing.

II.2a The Computer Model

The computer model used in this study has been under continuous development for

the past ten years and uses methanol as a fuel. During this,time, various

highlights of this model have been presented in the literature [11.2,3,4,5,6].

The model is a combination of two submodels, the kinetic Otto cycle submodel and

the kinetic exhaust submodel. The kinetic Otto cycle submodel includes both

thermodynamic and chemical kinetic considerations of processes occurring during

the intake, compression, combustion and expansion phases of the actual SI engine

cycle [II.2]. Modifications to the model to incorporate squish chamber cylinder

head geometry are outlined in Ref. 3. This model uses a detailed hightemperature chemical kinetic reaction mechanism [II.5] to define methanol

II.4
oxidation, and pollutant emissions formation of carbon monoxide (CO) and NO

during the combustion and expansion phases of the engine cycle.

The kinetics exhaust submodel is a one-dimensional fluid mechanics model of the

exhaust port and manifold which applies a detailed low-temperature kinetics

reaction mechanism [II.5] to predict unburned fuel (UBF) and aldehyde (ALD)

formation during the exhaust event [II.6],

Further refinements to the kinetic exhaust submodel stem from theories found in

recently published papers on the hydrocarbon quench layer [11.7,8,9]. These

studies suggest that under normal operating conditions, wall quenching is not an

important source of exhaust hydrocarbons. This is due to rapid diffusion and

burn-up of quenched hydrocarbons on a millisecond time scale without formation

of significant aldehydes [II.8]. These studies suggest that exhaust hydrocarbons

result from ring crevice storage and absorption-desorption of hydrocarbons in

the lubricating oil. Earlier studies by Wentworth have shown up to 74% reductions in.exhaust hydrocarbon concentrations by eliminating the ring crevice

[11.10]. Thus a submodel utilizing ring crevice storage was formulated to

calculate input hydrocarbon concentrations for the kinetic exhaust model. This

ring crevice storage model assumes that half of the unburned fuel existing in

the ring crevice at the end of combustion exits the crevice during the expansion

stroke and is subsequently burned in the bulk cylinder gases before the exhaust

valv e opens. It is also assumed that 10% of the unburned fuel in the ring

crevice does not exit the engine cylinder during the exhaust event. The

remaining 40% exits the cylinder under the following assumptions. One-half of

the exiting unburned fuel is evenly entrained in the bulk cylinder flow during

the time 90% of the cylinder charge (minus exhaust residual) is exhausted from

the cylinder. The remaining half exits the cylinder with the remaining 10% of

the bulk cylinder charge but at five times the concentration level. These

assumptions result in exiting hydrocarbon concentration profiles similar to

those found in experimental studies [11.11,12].

The method above was used in calculating input hydrocarbon level s for the

exhaust model in'al 1 but the exhaust gas recirculation (EGR) study. In this

part of the study, additional sources of unburned fuel emissions needed to be

considered. One such source resulted from reducing manifold vacuum as EGR was

II.5
added to maintain the same volumetric efficiency. This tends to increase fuel

flow through the cylinder during the valve overlap period due to scavenging.

Als o as EGR is added, there is a general degradation from good burns to slow

burns [11.13], thereby increasing unburned fuel emissions. Calculations using

only ring crevice volume resulted in decreasing input hydrocarbon concentrations

with increasing EGR. Thus for this part of the study, the input hydrocarbon

concentration was calculated for the no EGR case using the ring crevice model

and then used for all EGR addition cases as well.

Run times for the entire engine cycle simulation were typically 1100 seconds on

a CDC 7600 computer. This provided the best compromise between computer time

and numerical accuracy. A discussion of model tuning can be found in Ref. 5.

11.2b Spark Retard Effects

In a spark ignition engine, spark timing has a large influence on the combustion

process and, thus, on power, thermal efficiency and emissions. As shown in Fig.

.11.1, NOX emissions decrease rapidly as the spark is retarded from MBT (minimum

spark advance for best torque) due to burning later in the cycle and, thus, at

lower peak temperatures and pressures. The CO emissions tend to decrease with

spark retard due to higher temperatures during the expansion stroke. Unburned

fuel emissions also tend to decrease with spark retard due to higher exhaust

temperatures. Aldehydes, however, peak slightly retarded of MBT, then decrease

with more retard. While the increasing exhaust temperatures tend to react more

unburned fuel into aldehydes as the spark is retarded, the amount of unburned

fuel in the ring crevice decreases with spark retard due to lower peak

pressures. The later becomes the overriding effect for the spark slightly

retarded from MBT, thereby causing the aldehyde emissions to drop as the spark

is retarded further. Pischinger [11.14] found similiar results for lean

mixtures, but found a steady increase in aldehyde emissions with spark retard

for rich mixtures.

11. 2c Compression Ratio Effects

Due to the high octane number of methanol , higher compression ratios can be used

without engine knock. In conventional engines, squish chamber combustion

II.6
a

LU

f—<:

CJ

»—)

a

LU»—i

LJ

40

35

LU 30

0.20

o

CJ

Q_

Q_

0.10

0.00

200

ft 100

a

LU

a

«* 0

1000

a s:

LU a.

z o_

CD

500

LU

ID

§ = 1.00

9.00 CR

2000 RPM

MBT

^= 79.22

METHANOL

CO

ALD

12

a

10

 S g

9

3000

2500

2000

1500

1000

500

X

O

-20 -10 MBT 10 20

SPARK RETARD (DEGREES)

FIGURE I I.I: SPARK RETARD EFFECTS
geometry is used to minimize knock by reducing pressure rise rates at

the end of combustion and increasing surface area for heat transfer to occur.

Near the end of combustion the slower burning rates due to squish and the

increased heat transfer to the unburned gas minimizes end gas heating and thus

knock.

In this study, two methods of raising compression ratio are examined. The first

method maintains the same open-to-squish chamber volume ratio at top dead center

(TDC) while compression ratio is raised. This maintains the same relative

squish velocities with varying compression ratio and would have relatively the

same effect on power and emissions as raising the compression ratio in a flat

head engine. This method is best accompl ished.by piston crown redesign. The

second method maintains the same squish chamber clearance height while compression ratio is raised. Thus, compression ratio is increased by lowering only the

open chamber height. This results in increased squish velocities as compression

ratio is raised (in comparison to method one), thereby reducing the overall incylinder turbulence levels. This method is normally accomplished by milling the

cylinder head..

With the first method, NOX increased 29.6% by increasing the compression ratio

from 9:1 to 15:1 due to higher peak pressures and temperatures as shown in Fig.

11*2-* . tlnburhed fuel emissions rose 108.5% for the same comparison due to

increased ring crevice storage and decreased exhaust temperatures. Aldehyde

emissions, however, decreased 43.9% due to the lower exhaust temperatures,

thereby reacting less unburned fuel during the displacement flow period of the

exhaust; when aldehyde emissions are generally formed [II.5]. The CO emissions

rose l4*7%i ^ ;

.»"''*"'•' •

Using t-he_secbnd method, the increased turbulence levels and burning rate tended

to limft the increase in NOY and CO emissions as shown in Fig. II.3. By raising • . *"~ A *

the compression ratio from 9:1 to 15:1 by this method, NOX emissions rose 21.8%,

UBF emissions rose 109.8% and CO emissions rose 25.3%. Aldehyde emissions,

howeverj decreased only 19.3%., The trends of aldehyde emissions with compression ratio shown in Fig. II.3 are similar to the trends found by Pischinger

[IIsl4i using a similar method to raise compression ratio. Power and thermal

efficiency rose 15% for the same compression ratio increase.

11.8
I*

al

^ »-«

45

40

§ LU 35

i—i

a 20

~ a 10

O

a oo

^ 200

a.

o_

S3 100

a

LU

1 0

.1200

a -s.

LU a.

z: a.

m

600

LU

* = 1.00

9.00 CR

2000 RPM

MBT

°k= 79.2X

METHANOL

POWER

CO

NO

UBF

9 11 13 15

COMPRESSION RATIO

13

12 S

11

y3

10

3000

2500

2000

•1500-.

1000

500

X

O

FIGURE II.2: COMPRESSION RATIO EFFECTS

CONSTANT OPEN-TO-SQUISH VOLUME RATIO
LJ

LU

LJ

LU

40

§ LU 35

t—i

0.20

O

Q.

Q.

CO

LU

a

>-

LU

a

a 2:

LU a_

z a.

m _

LU

0.10

0.00

200

100

0

1200

600

0

* = 1.00

9.00 CR

2000 RPM

MBT

'V 79.21

METHANOL

POWER

CO

UBF

11 13 15

COMPRESSION RATIO

13

12

11

 2

10

-.3000

2500

2000

1500

1000

500

/•>

Q.

FIGURE I I. 3: COMPRESSION RATIO EFFECTS

CONSTANT SQUISH CLEARANCE HEIGHT
II.2d EGR Effects

In the EGR study, volumetric efficiency was maintained as EGR was added. A

throttled condition was chosen for this study to compensate for the increased

cylinder charge due to holding the volumetric efficiency constant while adding

EGR. Furthermore, the inlet charge temperature was allowed to rise with EGR

addition due to the mixing of the hot exhaust (600°K) with the cold incoming

charge (289°K). This raised the inlet temperature substantially from 298°K for

the no EGR case to 392°K for the 30% EGR addition at the 65% volumetric

efficiency condition. The increase in inlet temperature as EGR was added more

than compensated for the decreased burning vel ocities .thereby raising both

power and thermal efficiency as shown in Fig. II.4. Power and thermal

efficiency rose 8.5% with 30% EGR over the no EGR case, while NOV emissions A

dropped substantially showing a 97.2% reduction for the above comparison due to

lower peak cylinder temperatures and pressures. These results are consistent

with other investigators [11.1,15]. Unburned fuel emissions increased 55.8% for

the above comparison due to decreasing exhaust temperatures and oxygen

concentration. This matches trends found in our laboratory [II.1]. The EGR

effect reduced aldehyde emissions 88.3% and CO emissions dropped 70.7% for the

same comparison. Another interesting point was observed. As the exhaust

temperature decreases, more of the NO is converted to N02. For the no EGR case,

only 8.2% of the NOX is N02, while 30% EGR yields 49% of the NOX as N02.

It is evident that EGR is an effective method of reducing NO emissions without

A

large increases in UBF emissions.

11.2e Summary

A thermokinetic computer model was used to study the effects of simple NOx

control techniques on performance and exhaust emissions. A ring crevice storage

model was also formulated for predicting unburned fuel emissions. The study

showed spark retard to be an effective method of reducing NOX, CO and UBF

emissions at the expense of power, thermal efficiency and aldehyde emissions.

It was also shown that upon raising the compression ratio in a squish chamber

engine, the rise of NOX and CO emissions could be limited, but not without an

aldehyde penalty. Exhaust gas recirculation, however, was shown to be a very

11.11
effective method for reducing NOX. CO and aldehyde emissions with only a slight

UBF emission penalty. It appears from the studies that use of EGR in high

compression methanol fueled engines with increased turbulence to reduce cycleto-cycle variations will both maximize power and thermal efficiency while

minimizing harmful exhaust emissions.

11.12
a:

LU

LU

LU

45

40

LU 35

a 20

O

CJ

a.

o_

en

LU

a

>-

LU

a

a :E

LU a.

z a.

a: ^

m _j

z LU

a 10

0.00

200

100

0

1000

500

* = 1.00

9.00 CR

2000 RPM

MBT

1- 652

METHANOL

CO

10 20

EGR (%)

11

10

a QLU ^

9

 s

8

3000

2500

2000

1500

1000

500

X

O

30

FIGURE 11. 4: EGR EFFECTS
II.3 COLD START ENGINE PROCESS MODELING

Cold startability of an alcohol fueled vehicle is a vital issue since the

methods used to improve cold starting can affect warm-up driveabi1ity, emissions

and engine wear. As seen in Figs. II.5 and II.6, one of the cold starting

problems with alcohol fuels involves their low vapor pressures. If methanol is

inducted into a manifold at one atmosphere, the saturated vapor/air equivalence

ratio moves outside the lean flammability limit at a temperature of 12°C while

ethanol moves outside the lean flammability limit at 14°C. Under hard choking

conditions (0.67 atm manifold pressure) methanol moves outside the lean flammability limit at 5°C while ethanol moves outside the lean flammability limit at

8°C. Thus vaporization of alcohols in the intake manifold cannot be relied upon

for cold weather starting.

Current technology in alcohol fueled vehicles utilizes auxiliary starting aids

for enhancement of cold starting [11.16,17]. These aids are generally fuel

additives, auxiliary fuels, electric heat addition and dissociation or reformation of alcohol. Whil e all these methods will provide a cold start at low

ambient temperatures, there is a short period of time .following the cold start

in which intake manifold temperatures drop instead of rise and smooth operation

is difficult to obtain. It is this period which requires more than the one-shot

approach of auxiliary fuel devices-which can provide starting. Such systems

must be capable of supplying auxiliary fuel during part of the warm-up, but this

complicates the necessary cold start hardware. Electric heat addition will

assist in cold starting and warm-up in mild climates [II.3], however the amount

of electrical energy necessary to vaporize enough fuel for even idle conditions

can become prohibitive [11.17]. Dissociation or reformation of alcohols is a

very attractive method of providing cold starting enhancement since the hydrogen

gas it produces will allow starting at almost any temperature and the gas will

not recondense. However, large amounts of heat are necessary to dissociate the

alcohol and it is difficult to produce enough dissociated gases for the warm-up

period [II.3]. Finally alcohols can be tailored with such fuel additives as

butane, iso-pentane, gasoline and di-methyl-ether, to start at low ambient

temperatures, but these additives increase the cost, safety hazards and

emissions of the fuel. With these facts in mind, it can be concluded that the

ultimate solution to the cold starting problem has not yet been developed.

11.14
Ul

CJ

UJ

2.0 —

1.8 —

1.6 —

1.4 —

O

Ul

~ 1-0"

<:

\

cc

? 0.8 +

S 0.6-t

LEAN FLAMMABILITY LIMIT

-30 -10 0 10

TEMPERATURE (C)

20 30

FIGURE 11.5: SATURATED VAPOR/AIR EQUIVALENCE RATIO

VERSUS TEMPERATURE FOR METHANOL
2.0--

1.8--

1.6--

LU

CJ

g

C3

UJ

1..2--

K i.o

<:

\

o:

CD

CL

0

 0.8

S 0-6

QL

ZD

I—

<

CO

0.4--

0.2--

0.0

LEAN FLAMMABILITY LIMIT

-30 -20 -10 0 10

TEMPERATURE (C)

FIGURE 11. 6: SATURATED VAPOR/AIR EQUIVALENCE RATIO

VERSUS TEMPERATURE FOR ETHANOL
Part of the solution to the cold starting problem with alcohol fuels could be

more effective fuel transport and vaporization as they mix with the air. If

enough fuel can be inducted into the cylinder in a finely atomized fog, the work

of compression might vaporize enough fuel to provide a flammable,,mixture at the

end of the compression stroke [II.3]. The object of this study is to provide a

theoretical analysis of droplet transport and fuel evaporation ;in the., intake

manifold and engine cylinder during the intake and compression strokes and to

identify certain important design parameters affecting these phenomena, This

has been accomplished through a series of three separate models, the.Droplet

Fal 1-Out and Sling-Out Model, the Compression Process Model and the OnerDimensional, Two-Phase Cold Start Engine Model.

II.3a Droplet Fall-Out and Sling-Out Model

A simple fluid dynamics model, shown in schematic form in Fig. II.7, was formulated to study the effects of droplet size and air stream velocity on the

transport of a droplet along a two dimensional intake runOjer, :"T~hi:?;-.-_model

assumes that there is no droplet vaporization in the intake runne.rv, the air

stream velocity is constant and uniform throughout the runner (pTug fl.ow) and

the droplets are uniformly dispersed and do not interact. pimensifins for the

runner model are from a Ford 2.3 liter four cylinder engine intake .manifold.

"/ ' • ^" •

The droplet momentum equation is solved in Lagrangian form for both the x, and y

directions and can be found in Ref. 18. At the start of the'calculations, the

droplets enter the center of the runner with no initial velocity. The aerodynamic drag forces on the droplet from the moving air stream and gravity act on

the droplet as it moves along the runner. ""-.-- .

In Fig. II.8, an air stream velocity of 275 cm/sec was used to simulate a

cranking speed of 200 RPM. Droplet trajectories for various sized-.drppjets are

plotted within the intake runner. As can be seen from this figyrf, .the.lO and . . ~*.'*5p

i .,',- ^C""..*»^»vt

n,J,- " •, ' .'„

30 micron ( ) droplets will enter the cylinder while a 50 drpplet ea.ns|al".r put

of the air stream before entering the cylinder. The 100 droplet-is shown to

be slung out in the manifold corner. The larger droplets, having--lower surface

to volume ratios react more to gravitational than aerodynamic f&r.cjs and the

higher downward momentum during their travel in the down-draft sjjtettgn".of the

intake runner cannot be overcome. Shortening of this down-draft^seetiw^id not

11.17
INTAKE MANIFOLD DROPLET FALL-OUT AND SLING-OUT MODEL

*

/\\\v

/

/

*

;

^

/

///

///

/

\

>f

*/ FORCES ACTING ON DROPLET

/

/ Fdy

/

/

 dFx

 -J \JS

/

/ V 1

/ md g

/

^ y

FIGURE II.7: DROPLET FALL-OUT AND SLING-OUT MODEL SCHEMATIC
INTAKE MANIFOLD DROPLET FALL-OUT AND SLING-OUT MODEL

AIR STREAM VELOCITY 275 CM/SEC

/

/

/

/

/

/

',

\/

',

/

/

/

I

/, DROPLET SIZE

/ 1 / i nU

 j yU MM

/ — — 30 yuM

/ 50 ^

/, 100 ^H

/

/

/

/

</

A \^1 \ Vv ^V.

\\x -v^

\ \ ^

* 'w ~ ___ ___ ___ _^_ ____

\ ^^*—-

FIGURE 11.8: DROPLET SIZE EFFECTS ON DROPLET TRAJECTORY
have much effect on this problem as the larger droplets reach terminal velocities in very short distances.

Figure:'M.9 shows the effect of air stream velocity (or cranking speed) on the

trajectory of a 50 droplet moving in a horizontal plane. As can be seen from

this figure, the slower the air stream velocity (slower cranking speeds), the

larger the chance of the droplet entering the cylinder. This implies that the

centrifugal force is the the significant effect in droplet fall-out. For the

275 cm/sec case, the centrifugal force is over 25 times the gravitational force

indicating gravitational forces can be neglected in manifold design for this

droplet size and air speed range.

II.3b Compression Process Model

The compression process model examines the fate of droplets in the cylinder

during the compression process. The following assumptions are made to simplify

the analysis: (1) well stirred cylinder fluids, (2) all the fuel is in droplet

form at the beginning of the compression stroke, (3) an ideal mixture of air and

fuel vapor, (4) all droplets are of the same size and temperature, (5) all

droplets are in suspension during the compression stroke and (6) all cylinder

surface areas remain at a constant temperature.

The conservation equations are derived separately for the droplets and for the

gaseous mixture of fuel vapor and air and are listed in Ref. 18. Minimum engine

starting temperatures for methanol and ethanol are defined here to be the

temperatures at which the end of compression vapor/air equivalence ratio is at

the lean flammability limit (LFL) of 0.63 [11.19].

The compression process model was first used to examine the effects of fuel

vaporization on cylinder pressure and temperature. Using starting conditions of

298.15°K, manifold pressure of 0.775 atmospheres and a stoichiometric mixture of

50 liquid fuel droplets and air at the intake valve closing point of 60 degrees

after bottom dead center (ABDC), a compression ratio of 10 to 1 and a cranking

speed of 200 RPM, Fig. 11.10 shows a comparison of cylinder pressures during the

compression stroke for air only, air plus methanol, and air plus ethanol. As

can be seen from this figure, there is a substantial reduction in final cylinder

11.20
INTAKE MANIFOLD DROPLET FALL-OUT AND SLING-OUT MODEL

DROPLET SIZE 50 MICRONS

AIR VELOCITY

100 CM/SEC

— - 200 CM/SEC

— 300 CM/SEC

400 CM/SEC

FIGURE II. 9: CRANKING SPEED EFFECTS ON DROPLET TRAJECTORY
TJCD3Dm

o

CYLINDER PRESSURE (ATM)

ro

O)

GO

o

roO)

roo

r>

CDm3D

S 5-

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33a-oen

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s
pressure due to fuel vaporization in the air plus fuel cases. As expected,

methanol shows the greater reduction in final cranking pressure due to its high

latent heat and its larger fuel mass for a stoichiometric mixture in comparison

to ethanol.

Figure 11.11 shows the cylinder temperature versus crank angle for the same

comparison. As expected, methanol shows the greater reduction in final cylinder

temperature due again to its high heat of vaporization and increased fuel mass.

Fig 11.12 shows the fuel vapor/air equivalence ratio (lower curves) and droplet

size (upper curves) during the compression process for the above comparison. As

can be seen, both methanol and ethanol exceed the lean flammability limit

(vapor/air equivalence ratio of 0.63) at the end of compression stroke for room

temperature starting conditions. It also shows that ethanol droplets vaporize

more readily due to its lower heat of vaporization and decreased mass for the

same equivalence ratio. It should be noted that mass/ heat transport rates and

not vaporization saturation control the final values.

Fig. 11.13 shows the end of compression fuel -vapor/air equivalence ratio versus

starting temperature for methanol and ethanol holding all other starting

conditions constant. As can be seen from this graph, methanol has a minimum cold

starting temperature of 18°C for 50 drops while ethanol has a minimum cold

starting temperature of -12°C. Since methanol presents a much harder challenge

in cold starting based on this modeling study, the following study is limited to

methanol only.

Fig. 11.14 shows the effect of droplet size on minimum starting temperature. As

can be seen from this figure, droplet size plays an important role in the startability of an alcohol engine. As droplet size is decreased from 100 microns to

10 microns, the minimum starting temperature drops from 55°C to -15°C. A 30

droplet will al low starting above 3°C.

\

The effect of inducted liquid fuel/air equivalence ratio is shown in Fig. 11.15.

This figure shows that the amount of liquid fuel inducted into the engine

cylinder directly affects the slope of the end of compression vapor/air

equivalence ratio versus temperature curve by increasing the number of droplets,

hence, surface area for evaporation. The greater the amount of inducted liquid

11.23
700-r

650 --

600 --

550--

500--

450--

UJ

a

LU

a:

LQU_

UJ

o: 400

LU

o

350 '-

300--

250--

200

298.15 K

$ = 1.00

10: 1 CR

200 RPM

50 /uM DROPS

AIR ONLY

ETHANOL

— — METHANOL

0 20 40 60 80 100 120 140 160 180

CRANK ANGLE DEGREES (ABDC)

FIGURE 11.11: CYLINDER TEMPERATURE FOR ROOM

TEMPERATURE CRANKING
2.0-

1.8-

1.6-

o

1.4-

i— i

Ho:

B 1-2-

LJ

£ 1.0-

HD

o

LJAPOR/AI

R

pGO

=" 0.6-

0.4-

0.2-

n n

» •

298. 15 K

* = 1.00

10: 1 CR

200 RPM

50 iuM DROPS

ETHANOL

— — METHANOL

.__

"^^"^

"\^

\ \» -

\ y

Y

:

LEAN FLAMMABILITY LIMIT / \ /

/ /

/ ' \ // ' ' \

/

.<£

1 1 U- '" 1 1 1 1 1

r 80

- 70

- 60

-SOS

o

0£

CJ

t— 1

- 40 y

1— 1

en

COo

DROPLE

T

- 20

- 10

n

20 40 60 80 100 120 140 160 180

CRANK ANGLE DEGREES (ABDC)

FIGURE 11.12: FUEL EFFECTS FOR ROOM TEMPERATURE CRANKING
1.0

0.9

0.8

o

i—i

< 0.7

—»

a

0.6 —

0.5--.

0.4 —

LJ

g 0.3-

0.2--

0.1--

0.0

LFL

+

* = 1. 00

10:1 CR

200 RPM

50 HM DROPS

-30 -20 -10 0 10 20

AMBIENT TEMPERATURE (C)

FIGURE 11.13: FUEL EFFECTS ON MINIMUM

STARTING TEMPERATURE

30
1.0-r

UJ

METHANOL

10:1 CR

200 RPM

* = 1. 00

-20 -10 0 10 20

AMBIENT TEMPERATURE (C)

30

FIGURE 11.14: DROPLET SIZE EFFECTS ON MINIMUM

STARTING TEMPERATURE
1.0-r

0.9--

0.8--

0. 7 - -

o:

0.5-- ~

" a

u. ^ 0

 a: 0.4-

S<

LU

g 0.3-

0.2--

0.1--

0.0

LFL

METHANOL

10:1 CR

200 RPM

50 M DROPS

H 1

-30 -20 -10 0 10 20 30

AMBIENT TEMPERATURE (C)

FIGURE 11.15: INDUCTED EQUIVALENCE RATIO EFFECTS

ON MINIMUM STARTING TEMPERATURE
fuel, the steeper the slope of the curve. This trend however is offset by the

fact that more inducted fuel reduces the relative position of compression

vapor/air equivalence ratio curve as shown in the figure. Using the data shown

in the figure, the optimum inducted equivalence ratio would be approximately 1.5

which would result in a minimum starting temperature of approximately 16 °C.

In Fig. 11.16, the effect of cranking speed on droplet vaporization is shown.

As can be seen from this figure, the slower the cranking speed, the more time

available for droplet vaporization and therefore the lower the minimum starting

temperature. Decreasing cranking speed from 300 rpm to 100 rpm reduces the

minimum starting temperature from 24°C to 10°C for 50 micron methanol droplets.

In this cranking speed range wall heat transfer does not play a significant role

and of course cylinder leakage has been neglected.

Fig. 11.17 shows the effect of compression ratio on the startability of a

methanol mixture of 50 drops and air. This figure shows that small changes in

compression ratio have only slight effects on final vapor/air equivalence ratio.

However, raising the compression ratio from 10:1 to 20:1 enhances the minimum

starting temperature by dropping it from 18°C to -12°C for 50 methanol droplets.

In Fig. 11.18, the effect of manifold vacuum or choking is shown. Reducing the

starting cylinder pressure from 0.775 atm to 0.575 atm only reduces the minimum

starting temperature by 3°C. This can be anticipated when it is recognized that

fuel vapor pressure in all instances considered here is well below saturation

pressure.

II.3c One-Dimensional, Two-Phase Cold Start Engine Model

The one-dimensional, two-phase cold start engine model combines droplet transport and vaporization in the intake and exhaust manifolds with vaporization/

condensation of droplets during the compression and expansion processes of cold

start cranking. The following assumptions are made to simplify the analysis:

(1) a well stirred cylinder, (2) an ideal mixture of air and fuel vapor, (3) all

droplets are in suspension in the runners and cylinder and (4) all engine

surface areas are at constant temperature. In this model cylinder blow-by to

the crankcase is also considered.

11.29
LU

O

<->

METHANOL

lOjl CR

* = 1.00

50 M DROPS

-30 -20 -10 0 10 20

AMBIENT TEMPERATURE (C)

FIGURE 11.16: CRANKING SPEED EFFECTS ON MINIMUM

STARTING TEMPERATURE
1.0-r

METHANOL

* = 1.00

200 RPM

50 /uM DROPS

-30 -20 -10 0 10 20

AMBIENT TEMPERATURE (D

FIGURE 11.17: COMPRESSION RATIO EFFECTS ON MINIMUM

STARTING TEMPERATURE
METHANOL

* = 1.00

10:1 CR

200 RPM

50 uM DROPS

0.575 ATM

0.675 ATM

0.775 ATM

-30 -20 -10 0 10 20

AMBIENT TEMPERATURE (C)

30

FIGURE 11.18: MANIFOLD VACUUM EFFECTS ON MINIMUM

STARTING TEMPERATURE
As before the conservation equations are derived separately for droplets and the

gaseous mixture in the cylinder as well as all nodes within the intake and

exhaust runners. The engine is cranked for three consecutive cycles and the

fuel vapor/air equivalence ratio at the end of the cranking is examined for

f 1 ammabi 1 ity. A col or graphics output from this model is al so used to study

fuel transport and vaporization in the cylinder and runners during the cranking

process. Whil e this model is still under development, it has been shown to

produce reasonable trends when compared to actual engine data produced by the

experimental cold start engine (see Chapter III).

Initial study results from the model visually show the advantages of smaller

droplet size, electric heaters in the intake manifold, choking and dissociating

part of the liquid fuel for cold starting. Fig. 11.19 shows the vapor/air

equivalence ratio versus crank angle for three consecutive cranking cycles for

50 droplets injected into the air stream at an equivalence ratio of 2.0 (to

insure enough fuel enters the cylinder) for starting temperatures of 25°C and

0°C. The 10:1 compression ratio engine was cranked at 150 rpm for all runs.

Under these conditions with no manifold heat, the final vapor/air equivalence

ratio in the cylinder at the time of ignition (end of the third cycle) was 0.79

for the 25°C case and 0.57 for the 0°C case. Since a vapor/air equivalence

ratio of 0.63 is needed for ignition, a methanol engine under these conditions

could not be started by the third crank at 0°C. Decreasing the droplet size

injected to 30 increased the final vapor/air equivalence ratio to 0.70 at the

end of the third crank (flammable). Electric heat addition and placement were

al so found to have si gnificant effects. Placing a 200 watt electric heater

under the carburetor raised the final vapor/air equivalence ratio to 0.69 while

placing it in the intake port area raised it to 0.96 due to reduced heat

transfer to the cold intake manifold walls. Even at -20°C, with the heater in

the intake port area, the final vapor/air equivalence ratio was 0.78

(flammable). By choking at these conditions, the final vapor/air equivalence

ratio was raised to 0.89 at -20°C.

* Dissociating assumes that part of the methanol has been reformed into its

composing compounds: CH3OH -> CO + 2H2-

11.33
METHANOL

10:1 CR

150 RPM

* = 2.00

50 >jM DROPS

360 540 720 900 1080

CRANK ANGLE (degrees)

1260 1440 1620

FIGURE 11.19: STARTING TEMPERATURE EFFECTS
Dissociation of methanol is another effective way of increasing cold startability of an alcohol engine. If the amount of dissociated gas can be

minimized, this method can also be very attractive. By dissociating 15% of the

liquid fuel with a droplet size of 50 for the 0°C case, the final vapor/air

equivalence ratio was 0.65. Replacing 30% of the fuel with dissociated gases

raised the final vapor/air equivalence ratio to 0.72.

11.3d Summary

Three computer models were developed and are available for parametric study of

alcohol fuel transport, vaporization and ignitability in a cranking engine.

From the first model it is found that droplets must be smaller than 50 to

assure entering the cylinder under cranking conditions without being centrifuged

to the intake manifold wall. Since typical carburetors meter 100 to 500

droplets, cold starting without additional starting aids on alcohols becomes a

problem wel l above 0°C. This modeling study conclusion is consistent with

actual experience. Use of high pressure atomizing fuel injectors, however, can

atomize droplets and provide the fine mist of 50 droplets that can be carried

into the cyl inder during the intake stroke. Wei 1 nebul ized in-cyl inder fuel

injection is the ultimate in cold starting as the fuel does not have to travel

through a curved intake runner.

The second model faund that the heat of compression could be usedNto vaporize

enough fuel for starting under room temperature conditions. Ethanol was shown

to vaporize more readily than methanol due to ethanol's lower stoichiometric

mass and heat of vaporization. Stoichiometric ethanol mixtures of 50 micron

droplets and air were found to allow cold starting to -12°C while methanol under

similar conditions has a starting limit of 18°C. Droplet size was found to be

the most significant parameter in this study. A stoichiometric methanol mixture

of 10 micron droplets and air has a starting limit of -15°C. Very high compression ratios also aid cold starting showing marked reductions in cold starting

temperatures. All other variables, for the ranges examined by this study

(inducted equivalence ratio, manifold vacuum and cranking speed), showed very

little effect on cold startability of a methanol engine.

11.35
The third model showed that manifold heating, particularly in the intake port

area, was an effective means of increasing the cold startability of an alcohol

engine. Furthermore choking and small amounts of dissociated fuel were found to

aid cold startability. The color graphics output from this model was used to

study the details of the cold starting problem and solutions with much success.

It can be concluded from this study that fine droplet atomization and high

compression ratios will aid cold starting of alcohol engines and minimize the

amount of other cold starting aids. Furthermore, electric heat in the intake

port area greatly helps cold starting. The evidence produced here adds focus to

the historical question: is it more cost effective to convert a diesel by adding

a spark plug or a gasoline engine by adding injectors to produce an alcohol

engine or by blending volatile components into the methanol?

11.4 REFERENCES

11.1 R.K. Pefley, L.H. Browning, et al., "Research and Development of Neat

Alcohol Fuel Usage in Automobiles," Six Months Progress Report, NASALewis Contract No. NAG 3-143, University of Santa Clara Report No. ME81-2, July, 1981.

11.2 R.K. Pefley, L.H. Browning, et al., "Characterization and Research

Investigation of Methanol and Methyl Fuels," ERDA Contract No. EY-76-S02-1258, University of Santa Clara Report No ME-77-2, 1977.

11.3 R.K. Pefleyj L.H. Browning, et al., "Characterization and Research Investigation of Alcohol Fuels in Automobile Engines," Final Report, DOE

Contract No. DE-AC03-78CS, University of Santa Clara Report No. ME-81-1,

1981.

11.4 L.H. Browning, "A Thermokinetic Combustion Process Simulation for a

Methanol Fueled Spark Ignition Engine," Engineer's Degree Thesis,

Stanford University, Stanford, California, 1978.

11.36
11.5 L.H. Browning and L.A. Arganbright, "An Analytical Investigation of NOX

Control Techniques for Methanol Fueled Spark Ignition Engines," presented

at the Fifth International Symposium on Alcohol Fuels Technology,

Auk land, New Zealand, May 1982. ,

11.6 L.H. Browning and R.K. Pefley, "An Analytical Study of Aldehyde formation

During the Exhaust Stroke of a Methanol Fueled SI Engine," presented at

the Fourth International Symposium on Alcohol Fuels Technology, Guaruja,

Brazil, October 1980.

11.7 R. Bergner, H. Eberius and H. Pokorny, "Flame Quenching and Exhaust

Hydrocarbons in a Combustion Bomb as a Function of Pressure, Temperature

and Equivalence Ratio for Methanol and Other Alcohols," presented at the

Third International Symposium on Alcohol Fuels Technology, Asilomar,

California, May 1979.

11.8 C.K. Westbrook, A.A. Adamczyk and 6.A. Lavoie, "A Numerical Study of

Laminar Wal l Quenching," presented at the Western States Section

Combustion Institute Meeting, Berkeley, California, October 1979.

11.9 J.A. LoRusso, E.W. Kaiser and G.A. Lavoie, "Quench Layer Contribution to

Exhaust Hydrocarbons from a Spark-Ignited Engine," Combustion Science and

Technology, Vol. 25, pp 121-125, 1981.

11.10 J.T. Wentworth, "Effects of Combustion Chamber Shape and Spark Plug

Location on Exhaust Nitric Oxide and Hydrocarbon Emissions," SAE Paper

No. 740529, SAE Transactions, Vol 83, 1974.

11.11 W.A. Daniel and J.T. Wentworth, "Exhaust Gas Hydrocarbons —Genesis and

Exodus," SAE Paper No 686b, SAE Transactions, Vol 70, 1962.

11.12 R.J. Tabaczynski, J.B. Heywood and J.C. Keck, "Time Resolved Measurements

of hydrocarbon Mass Flowrate in the Exhaust of a Spark-Ignition Engine,"

SAE Paper No 720112, SAE Transactions, Vol 81, 1972.

11.37
11.13 W.G. Rado and W.J. Johnson, "Significance of Burn types, as Measured by

Using Spark Plugs as lonization Probes * with Respect to the Hydrocarbon

Emissions Levels in S.I. Engines," SAE Paper No. 750354, 1975.

11.14 F.F. Pischinger and K. Kramer, "The Influence of Engine Parameters on the

Aldehyde Emissions of a Methanol Operated Four-Stroke Otto Cycle Engine,"

presented at the Third International Symposium on Alcohol Fuels

Technology, Asilomar, California, May 1979.

11.15 J.A. Harrington, "Application of a New Combustion Analysis Method in the

Study of Alternate Fuel Combustion and Emission Characteristics"

presented at the Symposium on Future Automotive Fuels, General Motor

Research Laboratories, Warren, Michigan, October 1975.

11.16 L.H. Browning, J.F. Nebolon and R.K. Pefley, "Research Investigation of

Alcohol Usage in Spark Ignition Engines," SAE paper, presented at the

1982 ATD/CCM Conference, October 1982.

11.17 J. Nebolon, K.C. Chan, L.H. Browning and R.K. Pefley, "Multi-Point

Injection of Alcohol Vapor as a Cold Starting Aid for Neat Alcohol Fueled

Vehicles," presented at the Fifth International Symposium on Alcohol

Fuels Technology, Auk land, New Zealand, May 1982.

11.18 L.H. Browning, "Alcohol Cold Starting—A Theoretical Study," presented at

the 18th IECEC, August 1983.

11.19 T. Yano and K. It a, "Minimum Ignition Energies and Quenching Distances of

Methanol Blends," Bulletin of the Faculty of Engineering, Hokkaido

University, Japan, No. 94, 1979.

11.38
CHAPTER III

ENGINE COLD STARTING BY EXPERIMENTATION

In Chapter II cold starting is identified as one of the major problems impeding

the full use of pure alcohols as alternate automobile fuels. This problem has

attracted the attention of several research groups in the world yet their

efforts have met with only modest success with the pure alcohols below 5°C

[III.1,2,3,4].

11 I.I COLD START PROBLEM ELEMENTS

There are three aspects to the cold start problem. The first involves the low

vapor pressure of the pure alcohols , which limits the unassisted starting

temperatures under fully choked conditions to 12 °C and 14 °C, for methanol and

ethanol repectively, according to Section 11.3. Once the engine has started the

higher heats of vaporization of the alcohols and the higher flow rate of these

fuels cause significant temperature suppression in the intake manifold particularly in cold weather conditions near and below 0 °C . Under these conditions

it is reasonable to expect the fuel to condense on the intake manifold walls as

well as be present in both liquid droplet and vapor form. As vapor is needed

for ignition, this effect aggravates cold starting and idling. Additionally,

the suspended droplets and vapor when deposited between the cold spark plug

electrodes can act as an electrical conduit and prevent any further ignition,

making it very difficult to restart an alcohol fueled engine immediately after a

successful cold start if it fails to continue idling.

The second aspect of the cold start problem is related to the transportation of

the suspended fuel droplet through the intake manifold. The computer modeling

studies reported in Chapter II indicate that droplets above 50 microns may not

make it into the cylinder under normal cranking conditions. This is a very

important conclusion and it call s for better fuel preparation which is a

relatively unexplored area of the alcohol engine cold start solution strategy

[III.5]. Assuming that the droplets do make it to the cylinder, the third

aspect appears: What fraction of the liquid fuel can be vaporized during the

compression stroke, what is the fuel vapor to air ratio at the end of compression and what are the factors that adversely affect this vaporization. These

III.l
are the fundamental questions that have to been studied, both analytically and

experimentally, as prerequisites for designing and optimizing suitable cold

starting systems.

The third aspect of the cold start problem pertaining to the vaporized fuel

fraction has been an area of experimental and modeling study over the past

several years [III.I]. The amount of fuel vaporized during compression is the

important element in this study because the gaseous fuel component determines

the flammability of the air fuel mixture present at the time of ignition and the

task in the design of all cold start aids is to improve this number beyond the

lean flammability limit for all engine starting conditions. An efficient design

solution demands a fundamental understanding of the fuel vaporization process

and the experimental investigation conducted toward reaching this goal is

described in the following sections. The experimental approach taken here also

helps to validate and fine tune the computer model discussed in Chapter II.

111.2 EXPERIMENTAL APPROACH

During the compression stroke, the pressure and temperature of the gas inside

the cylinder increase. The high pressure forces some gas in the cylinder to

escape through the valves and rings (called blow-by) and the temperature elevation establishes some heat flow to the cylinder walls . These two factors

combine to produce cylinder pressure and temperature values less than predicted

for reversible adiabatic conditions. If the fluid in the cylinder is a mixture

of air and methanol fuel in fine droplet form, the pressure in the cylinder

tends to increase due to the addition of methanol vapor to the air but at the

same time the cylinder gas temperature tends to decrease due to the extraction

of heat from the surrounding air for the vaporization process. As alcohols are

volatile fuels with relatively high heats of vaporization, the pressure

reduction .due to temperature reduction predominates the pressure increase due to

the addition of vapor to the air resulting in a lesser magnitude of peak

cylinder pressure during compression of an air and fuel mixture compared to that

of air with no fuel (AIR-ONLY). The magnitude of this pressure suppression

depends on the mass of fuel vaporized during the compression process and this

quantity is influenced by the type of fuel used, the ambient air conditions, the

diameters of the fuel droplets and the droplet diameter distribution.

III. 2
The magnitude of gas leakage past the rings and valve s depends on the leak area

and the pressure difference hence, this process can be modeled as flow through

an orifice. The leak area can be determined experimentally as explained later.

The magnitude of heat transfer depends on the temperature gradient, the gas

Reynolds number and the shape of the combustion chamber (swirl etc). Although

much work has been done on the measurement and correlation of heat transfer

during cranking, none of the correlations can be considered universal. For this

reason it is necessary to experimentally determine the heat transfer and this is

best done by the analysis of the AIR-ONLY pressure data in which, once the leak

rate is determined, the only unknown quantity is the heat transfer to the walls .

By correcting for leakage and then determining heat transfer for the air only

case, the AIR & FUEL pressure data can be analyzed to obtain the rate of

vaporizaton and the mass of fuel vaporized during compression as functions of

crank angle. As a result the total mass of fuel available in the vapor state

and the air to fuel vapor ratio can be calculated at the time of ignition or at

the end of the stroke [111.6].

111.3 THE EXPERIMENTAL HARDWARE

The experimental engine for this study is the same 2.3 L Ford Pinto engine as

used in the Chapter II modeling study. It is mounted in a cold chamber test

cell (Fig. III.l). The engine head is suitably instrumented to receive a Quartz

crystal pressure transducer and a 0.005" diameter thermocouple (K-Type) in the

first cylinder. Figure III.l shows the experiment set-up and Fig. III.2 shows

the head instrumentation and the location of the pressure transducer. An

optical transducer that encodes each degree of crank rotation is coupled to the

crank shaft through a flexible coupling (Fig. III.3). Thermocouples are located

suitably to measure the temperatures of the intake manifold air, oil, coolant,

engine head, exhaust, and ambient air. A pressure tranducer is located at the

intake manifold to measure the manifold vacuum. These pressure and temperature

signals are suitably amplified and transmitted along with the pulses from the

shaft encoder through a high quality cable to the computer sampling ports.

A POP 11/03 computer is used in the data acquisition process (Fig. III.4). This

computer is programmed to record the gas pressure in the first cylinder for each

degree of crank rotation during a compression stroke, to sample the temperature

III.3
ORIGINAL PAGE 13

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FIGURE III.2: HEAD INSTRUMENTATION

LOCATION OF PRESSURE TRANSDUCER & THERMOCOUPLE

FIGURE I I I. 3: SHAFT ENCODER ASSEMBLY
COLD

CHAMBER

RE -AY h-TIMER

DELAY

ELIMINATION

CIRCUIT

CHARGE

AMPLIFIER

THERMOCOUPLE

AMPLIFIER

ENGINE

RPM CIRCUIT

MANIFOLD VACUUM

AMPLIFIER

INJECTOR

CNTRL CKT

-00-

ST1

DIGITAL

INPUT

DIGITAL

OUTPUT

A/D

POP 11/03

FIGURE III. 4 s SCHEMATIC OF EXPERIMENTAL SET-UP
and manifold pressure during the power stroke and, to send a control signal to

the fuel injector circuit before the start of a pre-selected compression stroke.

The control signal from the computer activates the injector circuit which is

designed to count a predetermined number of degrees after the engine bottom dead

center and activate the injection timing circuit. This timing circuit energizes

the injector (from a 12v DC source) and controls the duration of the fuel

injection; hence, the quantity of fuel injected during the compression stroke.

The counter circuit can count from 0 to a maximum of 60 degrees after bottom

dead center. The fuel injection system, an adapted Bosch cold start manifold

injector, is operated at 60 psi pressure to ensure good fuel atomization thereby

reducing the mass of fuel lost to the cylinder wall s (wall wetting). Since the

experiment is limited to cranking with no ignition, the injector is mounted in

the spark plug hole. The carburetor bowl is dried before the experiment and the

solenoid, mounted upstream of the carburetor, prevents any fuel flow into it

during the experiment. The shaft encoder is enclosed in an electric blanket to

maintain ambient operating conditions during cold temperature tests [111.6,7].

III.'4 THE DATA ACQUISITION SOFTWARE

The POP 11/03 micro computer has 16 A/0 ports, 16 digital read-in ports, 16

digital output ports, 2 external clocking ports (called ST1 & ST2 ) and a real

time programmable clock. The computer also has a host of library software

written in FORTRAN by the manufacturer for accessing the input/ output ports.

The flow chart of the data acquisition program is shown in Fig. III.5. During

the execution of the information routine, user defined names for data and text

files are created and some textual description of the experiment are accepted

interactively from the terminal and written to the text file. When the

instrument calibration routine is called, the computer samples the 16 A/D ports

continuously and displays the average values, standard deviation and 99%

confidence limit of the data on the terminal. The program executes this routine

until it receives an external interrupt on one of the digital input ports which

causes the program to jump to the main routine. Here, after servicing the

external circuits designed to eliminate data acquisition delay, the program

invokes the data acquisition routine and, after setting up the necessary

protocols for A/D operation, waits in a sampling mode to record the pressure

data during cranking. The data sampling routine is programmed to read, digitize

III. 7
INFORMATION ROUTINE

INSTRUMENT CALIBRATION

ROUTINE

HARDWARE SERVICE ROUTINE

SAMPLE 180 PRESSURE DATA ON

EXTERNAL SIGNAL (DEG. PULSE)

HARDWARE SERVICE ROUTINE

COUNT 330° CRANK ROTATION &

SERVICE EXTERNAL HARDWARE

ACTIVATE INJECTION

TIMER CIRCUIT

TAKE 180 PRESSURE DATA

READ OTHER CHANNELS

STORE DATA IN CORE MEMORY

STORE ON DISK AND STOP

FIGURE III.5_ : FLOW CHART FOR THE DATA ACQUISITION SOFTWARE
and store (1n computer memory) the pressure at the arrival of each degree pulse

In the ST1 port. The transfer of the degree pulse train to the ST1 port is

controlled by the delay elimination circuit which implements a sequential logic

operation based on the signal status of the bottom dead center pulse and the

control pulse from the computer digital output. After digitizing a total of 180

data points, control is transfered to the main routine in which the last three

pressure data points are averaged and tested for evidence of high pressure of

the compression stroke. This operation enables the computer to synchronize with

the engine. Once this is achieved, the cylinder pressure is sampled once every

degree during the compression stroke and the temperature, rpm and intake

manifold vacuum data are recorded during the fol 1 owing exhaust stroke. Just

prior to the preselected compression stroke the injector control circuit is

activated by generating a pulse at the digital output port for fuel injection.

After a user defined number of compression strokes have been sampled the data is

converted from binary representation to equivalent analog quantities and written

in the data file [111.6,8,9].

111.5 THE TEST MATRIX AND EXPERIMENTAL RESULTS

Two types of experiments were conducted: The AIR-ONLY tests in which no fuel

was injected and the AIR & FUEL tests in which a known quantity of methanol fuel

was injected for a selected stroke. These tests were conducted with various

ambient air conditions to study its effect on the mass of fuel vaporized. The

test matrix appears in Table 11 I.I.

TABLE III.I Test Matrix

TEST TYPE TEST TEMPERATURE °C FUEL USED MASS OF FUEL

AIR ONLY 23.0 5.0 0.0 -5.0 -15.0

AIR & FUEL 23.0 5.0 0.0 -5.0 -15.0 METHANOL 0.07 gr/cycle

Several tests planned for different stiochiometric ratios had to be canceled due

to the fuel flow rate limitation of the injector. At a speed of 100 RPM the

III.9
injection of 0.07 grams of fuel at 60 psi requires approximately 90 degrees of

crank rotatation and the cylinder pressure at the end point is in the range of

30 to 40 psi. At higher speeds or higher fuel amounts the cylinder pressure at

the end of injection will exceed fuel pressure, making further fuel injection

impossible.

Figure III.6 is a three dimensional plot of the AIR & FUEL test conducted at

23.0 °C. A stoichiometric amount of fuel was injected on the fifth cycle beginning at bottom dead center. The resulting pressure supression at the end of

compression is about 0.85 atm. Note that the subsequent peak pressures are

lower than before but are slowly rising. This is thought to be due to vaporization of the fuel sprayed on the wall or increased blow-by due to break down of

the oil seal. Figure III.7 shows the pressure suppression at different temperatures for the same mass of fuel sprayed. It must be noted that the A/F ratio is

stoichiometric only at 23.0 °C. At lower temperatures, due to the higher

density of the intake air, the A/F is greater than stoichiometric.

III.Ba Experimental Determination of Blow-by Area

The leak area is calculated from an experiment in which the crankshaft is locked

with the piston of cylinder #1 at top dead center. This cylinder is then

charged with pressurized nitrogen. After steady state conditions are attained,

the nitrogen supply is cut off and the resulting pressure and temperature

histories are recorded by the computer (Fig. III.8). The pressure and temperature decay data is used in a computer model to calculate the leak area.

The equations that adquately describe the engine blow-by are based on the

conservation of mass and the conservation of energy principals. The blow-by is

treated as compressible flow blow-down of a constant volume reservoir with heat

transfer. Since the cylinder boundaries are stationary and there is not much

bulk movement of the gas, natural convection heat transfer may be assumed. The

combined solution of these equations for an assumed leak area gives the temperature and density of the gas versus time from which the pressure can be computed.

The leak area and the convective heat transfer coefficient are modified until

the computed pressure and temperature conform to that obtained in the experiment. The leak area was computed to be 4xlO~3

cm2 [II1.6].

III.10
TEST DESCRIPTION B.B77 CR. FUEL INJECTED 5TH STROKE AT 68 PSI.

CYLINDER PRESSURE

VERSUS

CRANK ANGLE

RUN «i 198

TEMPi 25.0 C

CRi lltl

RPfc CRANKING

FUELf t€OH

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FIGURE III. 8 • PRESSURE CRANK ANGLE HISTORY.
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27 C

- - 0 C

15 C

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0 20 40 60 80 100 120 140

CRANK ANGLE DEGREES (ABDC)

160 180

FIGURE III. 7 PRESSURE DROP AT DIFFERENT

TEMPERATURES
0. 0 1.0 2. 0 3. 0 4. 0 5. 0 6. 0

TIME CIN SEC'S>

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FIGURE I I I. 8 a CYLINDER BLOW-BY.
III. 5b Determination of Valv e Discharge Coefficient

When calculations were first carried out, the valve lift was assumed to follow

the cam profile and this resulted in very erroneous values for heat flux,

especially in the region between 60 and 100 degrees of crank rotation (ABDC)

where the intake valv e opening changed from 0.5mm to 0.0mm. The valv e lift

profile was observed to have a pronounced effect on the heat flux calculations

because the duration of the intake valve opening determines the final magnitude

of the mass of air trapped in the cyl inder. This prompted di rect val ve 1 ift

measurements, which revealed significant departures from the lift derived from

the actual cam profile.he val ve assembly in the Ford Pinto engine houses a

hydraulic lash adjuster which in the absence of any oil pressure, absorbs part

of the cam displacement leading thereby to an early closure of the intake valve

during compression stroke as shown in Fig. 111.9. Since the experimental data

is taken during the first few cranks when the oil pressure is almost zero the

valv e lift profile can be assumed to follow the measured lift which changes from

0.05 mm at 60 degrees (ABDC) to 0.0 at 70 degrees (ABDC).

In order to determine the coefficient of discharge accurately at these low lifts

a flow bench was set up. The details of the experiment can be found in Refs.

III.2 and III.3 and the results are shown in Fig. III.10 [111.6,8,10].

III. 6 DATA REDUCTION AND ANALYSIS

The AIR-ONLY compression stroke is used to deduce the heat transfer. The

cylinder pressure of an AIR-ONLY compression stroke as explained earlier, is less

than the adiabatic pressure due to heat transfer and leakage. During the

initial part of the compression (between 60 to 90 degrees after BDC) the intake

valve is partially open and the total leak area is the sum of the constant leak

area determined experimentally and the variable leak area of the intake valve.

The intake valve area flow is computed from the valve lift and modified by the

discharge coefficient determined experimentally as a function of the valve lift.

The rate of loss of air from the cylinder to the surroundings or the rate of

mass addition from the intake to the cylinder for negative cylinder pressures is

computed based on the gas pressure, temperature and the leak area by so.lving the

appropriate gas dynamic equations.

III.14
1.0-r

0.9

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- rFROM CAM PROFILE

0 40 80 120 160 200 240 280 320 360

CRANK ANGLE DEGREES (ATDC)

FIGURE IXI.9i INTAKE VALVE LIFT DURING CRANKING
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INTAKE VALVE LIFT (IN MM)

FIGURE III. 10i INTAKE VALVE DISCHARGE COEFFICIENT
The general conservation equations applicable to AIR-ONLY tests are sol ved

simultaneously for a certain assumed heat transfer rate by a 6th order RUNGEKUTTA method to compute the pressure, temperature and density of the gas during

compression as functions of crank angle. The heat transfer rate is iterated

till the computed pressures and experimental pressures match. The heat transfer

computation was found to be extremely sensitive to even low noises in the

pressure data and so the data was curve fitted in sections with matching first

and second order derivitives at the common points. In the program, the step

size chosen was 0.1 degree of crank angle and the heat transfer computed at each

interval was stored in a temporary buffer and integrated to obtain the heat

transfer rate over a degree.

Figure III.11 shows the heat transfer computed by deduction in comparison with

the computer model of Chapter II, wherein the heat transfer is calculated from a

convective equation. Figure III.12 shows the computed cylinder gas temperature

i

in comparison with measured data. Figure III.13 shows the corresponding,

computed density in comparison with that calculated from the measured pressure

and temperature.

With all of the calibrations of the engine, instruments and data acquisition

system in place, the stage is now set for the calculation of the mass of fuel

vaporized from the pressure data of an AIR & FUEL test. Since the leak area and

the heat transfer rate are determined the general conservation equations

applicable to an AIR & FUEL test can be iteratively solved for an assumed

vaporization rate until both the predicted and the experimental pressures are in

good agreement [II 1.6,10,11,12,13].

111.7 MODEL VALIDATION

One of the goals of the experiment as noted earlier is the verification and fine

tuning of the cold start computer model (see chapter II). The computer model

was derived from first principles and uses a standard heat transfer correlation.

The leak area was based on the same experimental results as used here. The

model predicts the cylinder pressure for an AIR-ONLY and AIR & FUEL case with

ambient air conditions as a parameter. The mass of fuel vaporized is calculated

assuming a certain droplet diameter and size distribution. The validity and

III.17
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TEST TYPE

TEMPERATURE

ENGINE RPM

FUEL TYPE

THROTTLE

SYMBOL TYPE

AIR ONLY

27. O C

ISO. O

N/A

WIDE OPEN

- MODEL

o EXPERIMENT

0 20 40 60 80 100 120 140 160 180

CRANK ANGLE DEGREES (ABDC)

FIGURE III. lls HEAT FLUX

MODEL v« EXPERIMENT
650 T

600--

550--

500--

300

250

TEST TYPE

TEMPERATURE

ENGINE RPM

FUEL. TYPE

THROTTLE

SYMBOL TYPE

AIR ONLY

27. O C

ISO. O

N/A

WIDE OPEN

- MODEL

o EXPERIMENT

20 40 60 80 100 120 140

CRANK ANGLE DEGREES (ABDC)

160 180

FIGURE III.12 » CYLINDER TEMPERATURE

MODEL v« EXPERIMENT
16-r

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10--

LU

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FUEL TYPE

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SYMBOL TYPE

AIR ONLY

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ISO. O

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WIDE OPEN

- MODEL

o EXPERIMENT

4- 4-

20 40 60 80 100 120 140

CRANK ANGLE DEGREES (ABDC)

-I

160 180

FIGURE III.13 • CYLINDER PRESSURE

MODEL v« EXPERIMENT
accuracy of the model can be tested by comparing the experimental pressure data,

the computed heat transfer and mass of fuel vaporized with that of the model

prediction. Although exact agreement cannot be expected due to the differences

in the approaches, a good agreement in the magnitudes of the quantities involved

will go a long way in establishing a confidence in the model and improving the

accuracy of the parametric studies.

111.8 SUMMARY

To date the"model prediction and the experimental results are in close agreement

(Figs. III.11, 12, 13) [III.7]. Since the model uses a uniform droplet size of

50 microns the predicted magnitude of the pressure suppression is higher than

the experimental value where the droplet sizes are not only.much greater than 50

microns but are also distributed over a range of values. This difference points

in two directions: one, that current, poor cold starting of alcohol fueled

engines is in part due to insufficient nebul ization of the fuel and two, that

this discrepency must be overcome by using a better fuel injector at higher

pressure modifying the computer model to handle the droplet diameter distribution. The droplet distribution can be experimentally calculated from the laser

hologram studies and this approach is being currently explored. Although the

net heat transfer over the whole cycle is in close agreement (model and

experiment) it is felt that a better match can be obtained if the problem of

valv e lift is overcome by replacing the hydraulic lash adjuster by a solid

valve. During the follow-on work these issues will be sorted and the vaporization calculations will be concluded.

III.21
III.9, REFERENCES

1. R.K. j>efley, L.H. Browning, et al., "Characterization and Research Investigation of Alcohol Fuels in Automobile Engines," Final Report, DOE Contract

, . No., DE-AC03-78CS, 1981.

2. L.H. Browning, J.E. Nebolon and R.K. Pefley, "Research Investigation of

Alcohol Usage in Spark Ignition Engines," ATD/CCM Conference, Oct.1982.

3. J. Nebolon, L.H. Browning, et. al., "Multi-Point Injection of Alcohol Vapor

as a Cold Starting Aid for Neat Alcohol Fueled Vehicles" V International

Alcohol Fuels Technology Symposium, Vol. 2, 1982.

4. R.J. Nichols and R.J. Wineland, "Technique for Cold-Start of Alcohol Powered

Vehicles," V International Alcohol Fuel Technology Symposium," New Zealand

Vol 2, 1982.

5. L.H, Browning, "Alcohol Cold Starting—A Theoretical Study," 18 IECEC, Vol

2, Orlando, 1983.

6. G. Raghuraman, "Development of an Experimental Procedure 'to Calculate the

Mass Friction of Fuel Vaporized During the Cranking of an Alcohol Fueled

S.I. Engine," M.S. Thesis, University of Santa Clara, 1984.

7. L.H. Browning and G. Raghuraman, "A Detailed Study of Alcohol Cold

Starting," VI International Alcohol Fuels Technology Symposium, Ottawa,

Canada, 1984.

8. Microcomputer Handbook, Digital Equipment Corporation.

9. DECLab-03 FORTRAN Extensions User's Guide. Digital Equipment Corporation,

Maynard, Massachusettes.

10. G.P. Blair and J.M. Mcconnel , "Unsteady Gas Flow Through High-Specific

Output 4-Stroke Cycle Engines," SAE Paper 740736.

III.22
11. Yasuo Takagi, Yasuo Nakajima, Tsuneaki Saito, Tamotsu lijima, "Analysis of

Fuel Vaporization in S.I. Engine Cylinder During Low Temperature Cranking,"

SAE Paper 810 , 1981.

12. W.J.D. Annand, "Heat Transfer in the Cylinders of Reciprocating Internal

Combustion Engines," Proceedings of the Institute of Mechanical Engineers,

Vol. 177, No. 36, 1963.

13. A.C. Alkidas, "Heat Transfer Characteristics of a Spark Ignition Engine,"

Journal of Heat Transfer, Vol. 102, May 1980.

III.23
CHAPTER IV

VEHICLE AND ALLIED ENGINE STUDIES

During this grant there were two. major vehicle studies with neat alcohol. In

one study, supplementary funds were supplied by the United States Postal Service

to develop prototype kits for conversion of Ford Pinto vehicles to methanol or

ethanol operation. These kit designs were subsequently accepted and Alcohol

Energy Systems, under an independent project, supplied twenty of each to the

Postal Service where they were incorporated in the mail delivery fleets at

Denver, Colorado and Springfield, Illinois. In the other major study, the

Toyota Technical Center continued throughout this grant to make availabl e a

Toyota Cressida for evaluation of performance and emissions on methanol gasoline

blends.

IV.I. POST OFFICE PROJECT INTRODUCTION

The purpose of this task was to design, for the U.S. Postal Service, modification hardware to convert a small fleet of 40 Ford Pintos with 2.3 liter engines:

20 to methanol and 20 to ethanol, and maintain 49 state exhaust emissions

compliance. This was accomplished by engineering the required hardware and

creating a conversion manual which fully described the required hardware and the

procedures for the conversion [VI.1,2]. The hardware was assembled in the form

of kits and delivered to the Postal Service for their mechanics to install in

vehicles under an independent project. The manual is a unique and complete

description of the conversion of a specific vehicle class to the alcohols. To

aid in the development of the conversion specifications, two Postal Service

vehicles were loaned to the University. These vehicles are 49 state, 1980 Ford

Pintos, equipped with exhaust gas recirculation (EGR) and having CO, UBF two-way

(TW) catalyst systems.

IV.la Vehicle Modification Criteria

Four criteria were used in the development of the conversion specifications.

These criteria are: (1) performance and safety of the vehicle, (2) cost of the

modification, (3) fuel economy effects, (4) exhaust emission effects. The

development of the kit involved computer modeling, hardware development, and

IV.1
bench and dynomometer tests. The related computer models are described in more

detail in Chapter II, while the cold start hardware is described in Ref. IV. .

This section is primarily a summary of the actual conversion. The conversion

will be summarized under three categories: power and fueln economy, materials

compatabi1ity, and cold startability and driveabi1ity.

IV.la.l Power and Fuel Economy

To improve power and fuel economy, the compression ratio was increased from 9:1

to 11:1 by milling the cylinder head as the computer model predicted methanol

fuel economy improvements of 5% and 9%, for the postal service and urban

driving cycles, respectively (see Sec. II.1). For strength, the OEM head gasket

was replaced with the one used in the turbocharged version of the 2.3 liter

engine used in the Ford Mustang.

To provide the proper air fuel mixture for methanol or ethanol, eleven jets were

modified or changed in the Hoi ley 5200 carburetor. To improve acceleration

power, a vacuum delay valve was added to the EGR vacuum signal. This modification reduced the tendency to lean the mixture during acceleration. Thus, some

fuel economy was sacrificed for the sake of driveabi1ity.

IV.la.2 Materials Compatabi1ity

If not dealt with properly, materials compatabil ity problems can precipitate

system failures. The modifications used in this conversion were based on the

experience which is documented in the literature. It is common knowledge that

methanol will dissolve the terne plate coating of the fuel tank. The terne (90%

tin/10% lead) can clog the fuel filter and/or pass through the engine and

destroy the catalyst. To eliminate these problems, the terne coat was stripped

prior to operation on the alcohols. A woven polypropylene fiIter providing

eight micron absolute filtration was installed to prevent particulate fuel

contaminants from affecting the carburetor, engine, or catalyst system. The

carburetor was chromate dipped to inhibit corrosion, due the high electrical

conductivity of methanol.

IV. 2
IV.Ib Cold Startability and Drivability

Cold starting and warm-up driveability are significant challenges in the

development of alcohol conversion specifications for climates with temperatures

below 0 °C. The vehicles, to be put in service in Denver, CO and Springfield,

IL, required special attention both due to such low temperatures as well as high

altitudes at the Denver site when compared with the previous Pinto fleet

produced for the Department of energy [IV.1].

Many solutions to the cold start problem were considered. Fuel additives such as

iso-pentane and di-methyl ether were investigated. Considerable time and effort

were, spent on two cold start system prototypes. One system was a multipoint

vaporizer. The other was a gasoline injection system. This was developed and

then abandoned due to the complex feathering with the alcohol carburetor that is

required for multiple speed/load conditions during warm-up. The final choice

was the most practical, effective solution available.

An ether injection system, similar to those used in diesel engines, was used for

cold starting. Since the ether injection system operates during the cold start

only, other modifications were made to improve cold drivability. A barrium

titanate electric heater (180 watts) was installed under the carburetor to

improve fuel vaporization until heat from the engine was available. Such

heaters have temperature sensitive resistances which essentially create an open

circuit or. switch off at elevated material temperatures. The coolant flow rate

past the intake manifold was increased by a factor of thirty, and a copper

gasket was used between the head and intake manifold to increase heat transfer

to the manifold. The vacuum motor controlling the amount of heated air inducted

from the exhaust manifold was connected directly to manifold vacuum. This modification caused heated air to always be inducted, except at high loads, which

improved fuel preparation and driveability. Lastly, spark plugs one heat range

cooler were installed to prevent hot surface or pre-ignition, which can occur

with the alcohol s using standard heat range spark plugs under high load

conditions.

IV.3
IV.Ic Installation^of Kits and Field Experience .

The conversion kits were packaged in self-contained units, one per vehicle.

Prior to packaging., a Post Office mechanic visited the University of Santa Clara

where ,he studied the instruction sheet* watched the installation of one kit and

then performed the installation with a successful drive-away after the contents

of another kit were installed. The kits were subsequently shipped to the Postal

Service Centers .at Springfield, Illinois and Denver, Colorado where they were

installed, producing ten ethanol and ten methanol vehicles at each location.

The fueling sites at the Service Centers had been prepared with standard

metering pumps, fuel vapor recovery systems and filters, all of which were

certified alcohol resistent by the fuel system suppliers (see [IV.1]).

IV.lc.l Field Experience ,

The field experiences are being extensively documented in Ref. IV.3. Hence,

-only the highlights are touched upon here.

Although the field adaptation of the Pintos to a-lcohol went smoothly, in the

early phase of vehicle start-ups, several poor operational vehicle behaviors are

reported. These were traced to two circumstances: one was a carburetor jet

problem and the other was an installation problem..

The carburetor problem arose because the carburetors in the conversion kits were

not exact duplicates of those used on the prototype vehicles although the

carburetor manufacturer had originally certified them to be duplicates. This

required rejetting of the carburetors in the field. The second problem arose

because at one installation site the mechanics took their own initiative

relative to vacuum line hook-up rather than follow the specified instructions.

Another field problem arose with reports of excessive aldehyde levels around

some of the vehicles. This was traced to incorrect idle mixture adjustments,

plastic floats that had absorbed alcohol and, hence, en.richened the fuel-air

mixture and the type of Postal route driven. The mixture setting and float

problems were corrected by field servicing -of the vehicles which made aldehyde

levels acceptable. However, the park and walk route that some vehicles traveled

IV.4
would not allo w the hydrocarbon exhaust catalyst to reach full operating

temperature. As a result, the catalyst served as a very effective aldehyde

generator using the unburned fuel as an aldehyde source. This problem, while

not serious with a correctly tuned engine, deserves further attention if

aldehydes are to be minimized for all types of alcohol fueled vehicle usage.

A puzzling corrosion problem arose with the secondary throttle shaft of the

ethanol vehicles, while the methanol vehicles remained free of the problem.

This is contrary to expectation that corrosion problems are more severe with

methanol and ethanol. To understand the problem it is important to note that

exhaust gas recirculation (E6R) was required by the ethanol vehicles while it

was not required by the methanol vehicles to bring them into compliance with

EPA's 49-state emission standards. It has been observed (se« Chapter V) that

the condensate of the exhaust gases for alcohol fueled engines has a higher dew

point and may be more acidic than for gasoline vehicles. This combined with the

increased cooling of the carburetor from the high heat of vaporization of the

ethanol and the low power required for the postal driving route keeps the

secondary throttle shaft bathed in acidic condensate and corrosion results. The

absence of this condensate due to the absence of EGR in the methanol fueled

vehicles excluded the problem.

Engine corrosion and wear have been reported for the Pintos primarily relating

to the methanol fueled vehicles. The corrosion evidence is presented in Ref.

[IV.4]. Chapter V of this report gives a detailed explanation of the mechanism

as it is now understood.

IV.Id Summary

The Post Office methanol and ethanol fleet posed some unique challenges: very

cold tempertures, very low loads—primarily starting, driving short distances

and stopping.

The fleet was operated essentially complaint free as it relates to cold

starting. However, cold starting with a secondary fuel and fast idling to warm

engine state before driving are not considered to be fully satisfactory cold

weather protocol by the authors. The aldehyde problem and engine corrosion and

IV.5
wear evidence also deserve careful study in the interest of generating fully

successful field vehicles for the Postal Service.

IV.2 EGR STUDY

An engine dynomometer experiment was employed to study the effect of EGR on

emissions and fuel economy [IV.5]. There were two motivating factors for this

experiment. One, since EGR is used to decrease NOx emissions and noting that

these emissions are already lower for methanol than gasoline, could high EGR

rates reduce NOx emissions, possibly enabling the elimination of the two-way

catalyst (TWC). Two, since the aldehyde emissions are of major concern, could

the effect of high EGR rates be an effective way to reduce them.

A Ford 2.3 liter engine was used for this study, which was equipped with a

turbocharger, a feedback controlled carburetor, and a three-way catalyst system.

The EGR system was modified to provide manual control of the EGR rate. The spark

setting was also manually controlled and all data points were taken at MBT

settings. The emissions reported are NOx, UBF, and HCHO, measured by NDIR, FID,

and MBTH methods, respectively.

The EGR rate was varied over a matrix of test points: a fixed speed of 2000 RPM

at 1/3, 2/3 and full throttle, and a fixed load of 36 psi bmep at 1300, 2000,

and 3000 RPM. This load represented approximately 40, 35, and 20% of full load,

respectively. All test points were taken at stoichiometric air-fuel ratio and

MBT spark setting.

IV.2a EGR Limits

The limiting factor of allowable EGR is the lean misfire limit, which is

indicated by rough engine operation and high UBF emissions. This limit was found

to be approximately 25% EGR. A more practical limit is indicated to be 18% EGR,

at which the UBF emissions sharply increase.

IV.2b Variable Load Tests at Constant Speed

As can be expected, NOx emissions increased with increasing load, due to higher

IV.6
combustion temperatures. At high loads the NOx emissions could be reduced by a

factor of two, from 9.0 to 4.5 gm/kwh, with high EGR rates. At the low load

point, E&R reduced the NOx by an order of magnitude, from 8.0 to approximately

0.8 gm/kwh. Lower NOx rates with EGR are due to decreased temperatures from the

addition of the inert recircled exhaust gas. Unburned fuel emissions increased

with decreasing load and increasing EGR rate. Formaldehyde emissions followed

the same trends as the unburned fuel emissions, but were not as strongly

influenced by load or EGR rate as UBF. Brake specific fuel consumption improved

with increasing EGR due to two factors: decreased pumping losses as the throttle

is opened to compensate for the combustion mixture dilution and improved

mixture preparation due to the higher charge temperature.

IV.2c Variable Speed Tests at Constant Load

The NOx emissions increased with the increasing speed at constant brake man

effective pressure. High EGR reduced the NOx as much as 95% for these light

load points. Unburned fuel emissions increased with decreasing speed and, again

set a practical limit for EGR at about 18%. The HCHO emissions also increased

with decreasing speed, but were not strongly a function of EGR until 18% or

higher. As in the previous matrix, the BSFC improved slightly (about 5% average)

with higher EGR rates, until the misfire limit was reached.

IV.2d Conclusions ,

As in the case of gasoline fueled engines and consistent with the modeling

evidence in chap. II, EGR in the range of 0 to 18% is found to be effective in

reducing the NOx levels in alcohol fueled vehicles. The suppression can exceed

50% at high loads which suggests that lean burn engines which have high thermal

efficiencies may be successfully equipped with two-way instead of the three-way

catalyst systems currently being used on gasoline vehicles. Also, a gain in

fuel economy is possible by this technique. However, driveability may

deteriorate. Evaluations of these contrasting strategies need to be continued.

Aldehyde evaluations need also to be continued but the EGR effect on them

appears to be of secondary importance.

IV.7
IV.3 UNBURNED FUEL EMISSIONS CHARACTERIZATIONS

Unburned fuel emissions characterizations are important in the assessment of the

photochemical smog and direct exposure health effects of alternative fuels. For'

the alcohol s in particular, high hydrocarbon emissions may influence their

implementation. The high evaporative loss from gasoline blended with small

percentages (<10%) of methanol is a significant environmental factor retarding

the marketing of this fuel today. In the near future, utilization of neat OP

near neat alcohol fuels may be retarded due to concern over aldehyde emissions.

The objective of the study conducted at the University is to aid in assessing

the impact of these fuels, and als o aid the designer of alcohol systems to

design acceptable solutions.

Several segmental tests have been conducted for various reasons during this

contract period. The purpose of this article is to summarize the results of

these experiments(some, not previously published or reported) , and present them

in a meaningful way. Tables IV.1 and IV.2 are summaries of the exhaust and

evaporative emissions, respectively. This data is reported in the usual manner:

gm/mi for the FTP exhaust emissions, and gm/test for the FTP diurnal heat build

and hot soak evaporative emissions tests. Unfortunately, each test varied in the

kind of analysis performed. Therefore, if no detailed analysis was made, the

measurement is entered solely as UBF. If various components were identified,

these are listed, and their sum(except for aldehydes) is entered as UBF. These

data include the oxygen mass of the alcohol. The standard for exhaust emissions

is .41 gm/mi, while the evaporative emission standard is based on the sum of the

diurnal heat build and hot soak test, and is 2 gm. It should be noted that

emissions testing of alcohol and blend fueled vehicles can produce a variety of

answers depending upon the protocol (see [IV.6] and Sec. IV.4).

Figures IV.1 and IV.2 are summaries of the same data as in the tables, but in a

different format. To combine the results of the exhaust and evaporative tests,

the following assumptions made by the EPA were used [IV.7]. The gm/mi value for

the urban driving cycle is multiplied by1 29.4, the total vehicle miles per day.

To this value is added the diurnal heat build value plus 3.3 (the average number

of trips per day) times the hot soak value. This results in a total of grams of

unburned fuel emissions per day. The detailed emissions data is summarized in

IV.8
Table IV.1 Detailed UBF Exhaust Emissions Sunmary

VEHICLE

P.O. Pinto \

P.O. Pinto *

P.O. Pinto

 4

Fit I Pinto

Fit I Pinto

Fit II Rabbit

Fit II Rabbit

Fit II Rabbit

1982 Cressida

1982 Cressida

1982 Cressida

1980 Cressida

1980 Cressida

1980 Cressida

1980 Cressida

1980 Cressida

1980 Cressida

1980 Cressida

Cavalier

Cavalier

Celebrity

Celebrity

FUEL

M100

M100

M100

M100.

MEOH

M100.

MEOH

CDA-20

M80

M90

M100

Gas

M30

M70

M100

E30

E70

E100

Gas

M8.5

Gas

M8.5

UBF

8.20

6.72

.67

.31

.65

.66

.71

.47

.51

.48

.56

.29

.29

.44

1.19

.22

.30

.80

.15

.15

.13

.27

CH30H C2H50H

_ _

-

.23

.47

.63

.62

-

— _

_ _

-'

<.01 <.01

!05

.26

1.16

<.01 .03

<.01 .05

<.01 .65

_ —

-

-

CH4

_

- .

.05

.12

.03

.02

-

—

_

-•

.04

.04

.03

.01

.04

.04

.06

^ ,

-

-

NMHC C5M12

- -

-

.03 <.01

.06 .05

!07 !05

-

— —

_ _

-

.25

.20

.15

.02

.15

.20

.09

_ —

-

-

HCHO

.341

.097

.073

.052

.076

-

.110

.068

.075

.005

.004

.041

.090

_

.005

.001

.011

.122

-

C2H30H

_

-

- ' "

-

.273

—

_

-

.001

.^002

.001

<.001

_

.021

.066

_

-

-

5.5% Isopentane added

All emissions are expressed in gm/mi

Vehicle tested as received

Vehicle tested with carburetor float replaced

Vehicle tested with new float and jets replaced
Table IV.2 Detailed UBF Evaporative Emissions Summary*

VEHICLE

P.O. Pinto

Fit I Pinto

Fit I Pinto

Fit II Rabbit

Fit II Rabbit

Fit II Rabbit

1982 Cressida

1980 Cressida

1980 Cressida

1980 Cressida

1980 Cressida

1980 Cressida

1980 Cressida

1980 Cressida

Cavalier

Cavalier

Celebrity

Celebrity

FUEL

M100

MIOO^

MEOH

M100.

MEOH

CDA-20

, M90

Gas

M30

M70

M100

E30

E70

E100

Gas

M8.5

Gas

M8.5

UBF

-

4.65

4.07

1.26

1.90

3.28

1.42

1.06

2.76

4.51

-

2.27

2.34

1.32

.44

.44

.66

3.29

CH30H

-

4.65

3.27

1.26

1.23

-

-

•_

-

-

-

-

-

-

.00

.11

_

HC C6H6CxHy

-

.00

.80

.00

.67

-

_ .

_ _

-

-

_ .

-

-

-

.44 .21

.33 .19

_ _

C5H12 UBF

-

.00 1.06

.80 1.04

.00 .55

.67 .52

1.04

.39

.62

-

-

-

.70

_

-

.54

.55

.90

1.35

CH30H

-

1.06

.94

.55

.44

-

-

_

-

-

-

-

-

-

_

.08

_

HC C6H6CxHy C5H12

_

.00 - .00

.10 - .10

«- — mm

.08 - * .08

-

-

•- — «

_

-

- -

- -

- -

-

.54 .12

.46 .15

_

1 All emissions are expressed in gn/mi
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four basic groups of photochemical reactivity: aldehydes, non-methane hydrocarbons, alcohols , and methane. If no detailed analysis was performed, the

emission was assumed to be of the same composition as the original fuel. The

emissions contribution from the exhaust, hot soak, and diurnal heat build are

indicated by a "1","2", and "3", respectively.

IV.3a Description and Results of The Blend Tests

Two major studies and several experiments were conducted with blend fuels. One

study was the chacterization of the 1980 Toyota Cressida, which was conducted at

the Bartlesvilie Energy Technology Center [IV.8]. This was an exhaust emission

study measuring as many as 69 individual compounds. The fuels for the vehicle

were a gasoline baseline, M30*. M70, and M100, E30, E70, and E100. In addition,

several SHED tests with these fuels were conducted at the University; however,

no gas chromatograph analysis was performed.

The values obtained from these tests can be found in the tables; the trends will

be discussed here. With the oxygen mass included, total UBF emissions were

comparable to the gasoline baseline for the M3Q, E30, and E70 tests, and the .41

standard was met for all these fuels. This statement could also be made for the

M70 tests, with the oxygen excluded. Only on pure methanol and ethanol were the

UBF emission rates significantly higher. The aldehyde rates were significantly

more than the baseline for the 70 and 100% runs, and were consistently about 8%

of the UBF emissions. Interestingly, the evaporative emissions contribution is

as much or more than that of the exhaust emissions for some of these blend

fuels. None of the blend fuels tested could meet the evaporative standard, even

with the oxygen mass excluded. Note that this was basically a gasoline vehicle,

and was not optimized to run on neat alcohol, nor high blend percentages.

Another study was conducted at the University with two 1983 vehicles. The fuels

tested were gasoline and a post refinery blend of 90% unleaded gasoline, 8.6%

methanol and 1.4% iso-butanol as a cosol vent. Two GM vehicles were used: a

Cavalier, which was fuel injected, and a Celebrity, which was carbureted. The

30% methanol by volume added by post refinery mixing to the gasoline.

IV.13
interest was primarily in the evaporative emissions effects. The methanol and

aromatic components of the hot soak and diurnal heat build emissions were

measured; no such analysis was made of the exhaust UBF measurements.

Surprisingly, the fuel injected vehicle showed no increase in evaporative

emissions, even with the oxygen mass incl uded,, and met the standard on both

fuels. Exhaust emissions decreased by 3% with the blend. The carbureted vehicle

showed a three-fold increase in evaporative emissions on the blend, and failed

to meet the standard. The exhaust UBF emissions doubled, but remained within

the .41 gm/mi standard.

To summarize the results of this study, the short term evaporative emission

changes with a 10% oxygenate blend were excessive for a carbureted vehicle, and

negligible for a^ fuel injected vehicle.

IV.3b Description and Results of the Neat Fuel Studies

One major study and many experiments were conducted involving the measurement of

UBF and aldehydes as well as a detailed analysis of the UBF emissions with neat

or near neat alcohol fuels.

A series of tests was conducted on an in-service Post Office vehicle operating

on methanol. As reported in the previous section, the Post Office vehicles on

occasion were operating extremely rich, and this series of tests was meant to

investigate the causes and resulting exhaust emissions effects of the rich

operation. Based on CO'measurements, a test vehicle was found to have an

equivalence ratio at idle of approximately 1.3. The vehicle was tested as it was

received. The CO emissions were extremely high (100 gm/mi) and the UBF emissions

were 8.20 gm/mi. Aldehyde emissions were also extremely high (0.341 gm/mi). It

was first thought that the rich operation was due to the carburetor float

sinking from absorbing' methanol. This cause was first tested by replacing the

float and retesting the vehicle. The UBF emissions dropped to 6.72 gm/mi and

the aldehyde rate dropped to 0.097 gm/mi. Then, the fuel and air jets were

replaced with their proper sizes, and the emissions were markedly decreased

further: UBF, 0.67 gm/mi; and aldehydes, 0.073 gm/mi. The conclusion of this

study as described in Sec. IV.lb was that a mi seal ibrated alcohol car can be a

IV.14
major aldehyde emitter, and , as expected, a heavy carburetor float can produce

a measureable difference in emissions.

The major characterization with neat alcohol fuels was of the California Energy

Commission Fleet I and II vehicles: 1980 Ford Pintos and 1981 Volkswagen

Rabbits, respectively [IV.9,10]. The fuels used were neat methanol , and a 5.5%

mass iso-pentane/methanol blend. The vehicles were tested for UBF emissions of

methanol and Cj to Cg hydrocarbons. The aldehydes reported here were taken from

another series of tests which have been presented in a previous progress report

and elsewhere [IV.9,10]. A Fleet II Rabbit on denatured ethanol was also

tested (without GC analysis), and the results are reported here [VI.10].

The UBF emissions were less for both Fleet I and II vehicles on methanol when

compared with the isopentane blend. The Pinto vehicle met the 0.41 gm/mi

standard with the oxygen mass included; all other cases met the standard with

the oxygen excluded. The Pinto UBF exhaust emissions averaged 17% by mass

methane, compared to 3.3% for the Rabbit. The evaporative emissions were quite

high for the Pinto, due to that fact that it is carbureted, and that the

cannister was probably heavily loaded with methanol after 13,000 miles of

operation. The Rabbit met the 2 gm evaporative standard on methanol with the

oxygen included, and on the isopentane/methanol blend with the oxygen excluded.

Aldehyde emissions were measured for both cars only on their "standard" fuels

(neat methanol for the Pinto and the methanol/isopentane blend for the Rabbit).

Formaldehyde rates were 16% of the UBF rates for the Pinto and 11% of the UBF

rates for the Rabbit.

With the oxygen mass excluded, the Rabbit on denatured ethanol met the exhaust

standard, but failed the evaporative standard. The aldehydes level, using the

MBTH technique, and calculated as acetaldehyde, was .273 gm/mi, which was 58% of

the UBF emission.

A 1983 Toyota Cressida (see Section IV.4) was also tested for aldehydes and

evaporative emissions. The vehicle was tested on M80, M90, and M100 and various

spark settings (0°, 2° and 4° BTDC) were also used. The UBF results were very

CDA-20 denatured with 10% of gasoline type hydrocarbons.

IV.15
similar for all three fuels;' The average aldehyde rate for the three fuels was

0.084 gm/mi, about 15% of the UBF emissions were measured for the M90 blend and

were quite low: 1.81 gm total with the oxygen included.

IV.3c Concluding Observations

Several observations can be drawn as a summary of these studies:

1. Short term evaporative emission effects can be excessive with post refinery

blend fuels, particularly with carbureted vehicles.

2. Long term effects of evaporative emissions on neat fuels should be investigated further, due to the lowered purging efficiencies of the cannister.

3. A properly tuned vehicle optimized for alcohol will produce acceptable UBF

and aldehyde emission rates, and may even be superior to gasoline.

4. Fuel injection systms may be superior to carburetion systems in adapting to

small fuel composition shifts; hence, in controlling exhaust and evaporative

emissions.

IV.4 TECHNIQUES FOR EVALUATION OF EMISSIONS OF TESTING ALCOHOL FUELED VEHICLES

The Federal Test Procedure for light duty vehicles defines the method for

determining emission rates and fuel economy for comparison with the regulated

standards. This test procedure is based on gasoline or diesel, and to maintain

data integrity, must be modified for alcohol. Many of the modifications were

described elsewhere [IV.7,11] and are summarized here. The primary areas of

concern are unburned fuel emissions, fuel economy, and exhaust condensation.

IV.4a Unburned Fuel Emissions

For gasoline and diesel, a H/C ratio of the emissions is assumed for the exhaust

emissions, hot soak, and diurnal heat build tests. A similar assumption is

required whenever an alternative fuel, such as alcohol, is used. The accuracy of

the assumption will limit the accuracy of the mass emission measurement. Based

on the tests of the 1980 Cressida and the Fleet I and II vehicles previously

described, the assumption that the UBF emissions are of the same composition as

the original fuel was checked. This simple assumption resulted in errors of 2 to

IV.16
26%, primarily due to a decrease in the oxygenate; fraction in the emission

relative to the fuel.

It is important to reach a common agreement as to how the UBF emission is to be

defined,i.e. whether or not the oxygen mass is included as it often determines

whether or not the vehicle meets the standard. From photochemical considerations (see Chapter VI and [IV. 11]), which are the basis for control 1 ing UBF

emissions, it is reasonable to define the emission without the oxygen mass

associated with the alcohol molecule. Excluding the oxygen mass results in a

secondary benefit: the accuracy of the UBF emissions measurement improves. Based

on the same tests, but now excluding the oxygen, the assumption that the

emission is of the same composition as the fuel results in errors of only 0 to

2.4%.

The -measurement of unburned fuel emissions is further complicated because the

response of the flame ionization detector (FID) to alcohol is reduced by 10 to

30% compared to propane, the standard calibration gas. This response factor

varies with the instrument model, burner configuration and operating conditions.

A technique was developed at the University to determine the response factors of

its instruments [IV.7]. Further experiments showed that the response factor is a

linear function of the components of a blend sample. Thus the best solution at

this point is to determine the response factor for a particular FID for methanol

and ethanol , and calculate the response factor for any blend fuel tested. The

assumption of a response factor of unity for the hydrocarbons would be

consistent. with current FTP protocol.

IV.4b Fuel Economy Calculations

The calculation of fuel economy for alcohol fuels is incorrect using the carbon

density of gasoline. The errors for methanol and ethanol are of the order of 116

and 56% high for the neat fuels, respectively. Using the correct carbon density

of the fuel will result in a correct volumetric fuel economy. However, a

volumetric based fuel economy is not the best way to compare fuels of different

energy densities. An energy based fuel economy, such as mi/106 Btu is more

appropriate.

IV.17
IV.4c Water Condensation •••"*! '

Stoichiometric air/fuel mixtures produce a methanol exhaust which contains 23

molar percent water, an ethanol exhaust which contains 18 molar percent water,

compared to 12 percent for gasoline. If corrective action is not taken,

condensation will occur in most exhaust dilution tunnels used in CVS testing

when evaluating these alternate fuels. Condensation wi 1 1 cause UBF and NOX

emissions to be'artificial ly low. Calculations have been performed that

determine what temperature is needed to avoid condensing the exhaust mixture for

a given vehicle and blower flowrate. For the system at the University,

condensation is prevented by heating the dilution air to 120 F, which has proven

adequate. Condensation has not been observed in the sample bags. This may occur

but can be avoided by recognizing the difference between the exhaust condensate

of gasoline and alcohol fuels.

IV.4d Conclusion

The aforementioned areas of concern should always be addressed whenever one is

testing an alcohol fueled vehicle. The issues raised should be the subject of

discussion among testing groups, leading to the establishment of a defined test

fuel, a formaldehyde emission standard, and a test procedure. This achievement

is seen by the authors as a necessary step to the development and commercialization of alcohol fuels.

IV.5 TOYOTA CRESSIOA—A UNIQUE TEST VEHICLE

IV.5a Introduction

Having recognized the alcohol fuels as extenders and replacers of petroleum

based fuels in the transportation industry, in the near term it is possible to

accommodate the use of '5% anhydrous methanol or 10% ethanol as gasoline

extenders for current production vehicles as is evident by Environmental

Protection Agency waiver s for such fuels and earlier work [IV.13]. Although

current alcohol production may be used in blends, it is presently nowhere near

the level required to provide fuel for a significant number of pure alcoholfueled vehicles. Conversely, manufacturers are reluctant to produce vehicles

IV.18
capable of operating on alcohol until an abundant supply of fuel becomes

available. One solution to this circular impediment to utilization lies in the

development of an inexpensive modification to an existing vehicular engine

control system to permit operation on any blend of gasoline and alcohol. If a

large number of such vehicles were available, implementation of alcohol fuels

could be carried out in a more flexible and expeditious manner.

Multipoint Electronic Fuel Injection systems (EFI) with feedback control are

currently, appearing in new cars and have excellent potential for meeting the

above requirements as well as maintaining fuel economy and emissions levels.

The feedback systems provide precise air-fuel (A/F) ratio control and thus offer

automatic compensation for the shift in stoichiometry required by the alcohol

fuels thereby retaining performance, emissions, fuel economy, and driveability.

They are also attractive for use with alcohol fuels if properly compensated for

fuel stoichiometry shifts since they inherently provide better acceleration

transient response, warm-up driveabi1ity, and essentially no maldistribution of

the air-fuel mixture among the cylinders; characteristics which have been

persistent problem areas for carburetted vehicles and are exacerbated by the

alcohols [IV.14]. When EFI is used in conjunction with a three-way catalytic

exhaust system, very low levels of exhaust emissions can be obtained. Correct

combustion stoichiometry required for optimum control of exhaust emissions in

such three-way-catalyst (TWC) systems requires some type of feedback logic.

This feature can automatically maintain stoichiometry for al cohol/gasoline

blends to some blend level depending upon the compensation range of the EFI

system. The Toyota system has been shown to have a wide inherent compensation

range [IV.15].

IV.5b The Toyota Closed Loop Feedback Controlled Fuel Injection System

The Toyota EFI system represents the "state-of-the-art" in current EFI systems.

It has been adapted from the Bosch L-Jetronic system. A schematic diagram of

the fuel system components is presented in Fig. IV.3.

An air flow meter provides the major input to the electronic control unit (ECU).

The exhaust oxygen sensor "feeds back" information about combustion product

composition so that the ECU can fine tune its fuel injection duration commands.

IV.19
EXHAUST

f

E.G.R

FIGURE IV. 3 TOYOTA EFI CLOSED LOOP CONTROL SYSTEM
All injectors are connected electrically in parallel and simultaneously inject

one-half of the required fuel to each cylinder once each revolution of the

crankshaft. The resulting air-fuel ratio alternates rapidly about the stoichiometric mixture within a narrow range. The TWC system requires such a precisely

controlled mixture in order to optimally reduce nitric oxide (NOX) while simultaneously oxidizing carbon monoxide (CO) and unburned fuel (UBF) emissions.

Reference IV.15 presents evidence that an unmodified Toyota EFI system with

feedback control can accommodate a post refinery blend of 50% ethanol/gasoline

(by volume) or 30% methanol/gasoline before the leaning effects cannot be fully

compensated and the leaning effect on emissions limits further increase in

alcohol percentage. These blends represent a stoichiometric compensation of 18%

relative to gasoline. The ECU of a 1980 Cressida was subsequently modified in a

minimal fashion so that the impulse time generated by the control multivibrator

could be selected manually in ten steps ranging from 1.0 to 2.0 times the

original injection duration. The oxygen sensor coupled with the feedback logic

was relied upon to maintain stoichiometric conditions for all fuel blends

tested. Pre-test checks were made to assure that the system was within the

authority range of the feedback control system for each blend. A supplemental

10 micron polypropylene fuel filter was also installed downstream of the

original filter to keep the EFI system free of any particulates loosened by the

presence of the alcohols blended into the gasoline.

These changes al 1 owed the expl oration of the concept of a "smart car", i.e., a

vehicle capable of handling any combination of alcohol and gasoline.

IV.5c Test Results and Discussion--1980 Cressida

Using the 1975 Federal Test Procedure (FTP), the vehicle was evaluated over the

entire spectrum of gasoline/alcohol blends for the purpose of evaluating driveability, fuel economy, and emissions. Dilute exhaust emissions were measured

using conventional instrumentation and FID response factors were applied for the

methanol and ethanol fractions of the exhaust [IV.10]. Evaporative emissions

were quantified as functions of blend percentage. Detailed unburned fuel

characterizations were conducted for the purpose of having an inventory useful

in photochemical reactivity studies. Aldehyde emissions were also studied.

IV.21
Carbon monoxide emissions, meeting for all blends the 1980 California emission

standard of 9.0 gm/mile, were generally unchanged during the alcohol blend test

series and the fuel economy on an energy basis remained constant regardless of

the blend used [IV.15]. Good driveability was maintained for blends ranging

from pure gasoline to 90% ethanol and methanol.* This suggests that the

injection duration modifications incorporated with the electronic control logic

maintained near-stoichiometric operation with alcohol blends in all driving

modes, otherwise a substantial change of CO emissions and mileage, and/or

serious driveability problems would have been observed.

Hydrocarbon emissions calculated as UBF, show very consistent emission levels

for light and moderate blends of both alcohols. For higher percentage blends,

the UBF nearly doubled when compared to gasoline. Part of this increase can be

attributed to the oxygen mass contribution of the alcohol. On a carbon basis,

the unburned fuel remained constant or decreased for methanol blends of 50% or

greater, and all ethanol blends. For the lower blend percentages of methanol,

carbon-based unburned fuel emissions increased by 24 to 55% over the Indolene

baseline. In all instances, the blend HC emissions were below the 1980

California standards of 0.41 gm/mile [IV.15].

The post catalyst NOX emissions were higher for the blend fuels and generally

increased with alcohol percentage. Lower engine exhaust NOX emissions for neat

alcohol fuels relative to gasoline are repeatedly reported and it is true here.

The implication is decreased TWC efficiency. There is other evidence of this

rising NOX with other TWC fuel injected vehicles operating on blends suggesting

a generic problem [IV.10].

IV.5d Rise of N0y Emissions From Alcohol Blends

To investigate the NOX increase with increasing alcohol blend, attention was

focused on the oxygen' sensor [IV.16,18]. Dummy catalyst (no catalytic

materials) tests confirmed engine-out NOX emissions were lower for the alcohols

than with Indolene, supporting

The 90% ethanol fuel contained 10% hydrocarbons as a denaturant and the

90% methanol contained 10% gasoline for cold startability.

IV.22
prior research. Therefore catalyst performance must be affected by the addition

of alcohol and/or transient operation factors. A lean bias of the 02 sensor

associated with exhaust temperature increase during acceleration for alcohol

fuels was predicted using a form of the Nernst equation. A lean mixture can

result in decreased NOX conversion efficiency for a TWC. Therefore, if during a

vehicle acceleration when most of the NOX is formed and exhaust temperatures

increase, the mixture shifts lean and the oxygen-rich exhaust lowers NOX

catalyst efficiency, the result could be a relative increase of post-catalyst

NOX emissions in comparison with Indolene.

To evaluate this possible lean bias, several modifications were made. First,

the 02 sensor reference voltage in the feedback loop was raised from 0.45 to

0.75 volts in 0.05 increments and the integrator output which divides the cycle

into rich and lean periods was changed to a non-symmetric profile. Second, an

oxygen sensor composed of a thinner zirconia ceramic was installed, resulting in

a 20% faster air-fuel ratio modulation. The Cressida was then tested using

various combinations of these modifications.

The effect of the feedback control modifications was a significantly richer airfuel mixture. Tailpipe CO and UBF emissions increased with slight reductions in

NOX for all fuels. However, NOX emission levels for alcohol were still above

those for the baseline Indolene. Addition of the fast response oxygen sensor

produced a slight leaning effect which was evident from decreasing CO and UBF

and increasing NOX. This was opposite to the intended effect.

In conclusion, the overall effect of the modifications showed that improvement

in engine-out emissions was possible through improved stoichiometry management,

but the TWC efficiency would still need improvement if the addition of alcohol

fuels is to result in an NOX reduction relative to gasoline as expected.

IV.5e Operating Experience

The Cressida was extensively operated on all blends of alcohol and gasoline for

45,000 miles. Several material compatibility problems arose. The terne plate

coating of the fuel tank was removed by the methanol. Due to some vapor lock

problems the fuel filter had been removed and the terne plate particles plugged

IV.23
the fuel injection system. The coating was subsequently removed chemically and

the injectors replaced. The original fuel immersed electric fuel pump failed

due to galvanic corrosion. It was replaced with a corrosion resistant pump

which had its electric drive isolated from the fuel.

Vapor lock was experienced on a 10% methanol blend at high altitude and warm

weather. It was caused by improper placement of the additional fuel filter.

After relocation of the filter, no additional problems occurred.

At approximately 43,600 miles, abnormal engine noise became apparent. Upon

disassembly it was found that three out of the six connecting rod bearings had

spun around the crankshaft throw. Visual inspection showed evidence of chemical

attack. This was another piece in the body of growing concern for corrosion

and wear caused by the alcohol fuels. (See Chapter V, Corrosion and Wear

Mechanisms). The engine was subsequently rebuilt and limited testing performed.

IV.5f Test Results and Discussion--1982 Cressida

A 1982 Cressida was provided by the Toyota Corporation to continue the study of

the NOX issue. It subsequently was equipped with a high compression ratio

engine and tested on various methanol/gasol ine blends. It was supplied as a

gasoline blend vehicle, i.e., the fuel system parts were alcohol compatible, the

distributor modified as well as the ECU. The high compression ratio engine has

a CR of 12.5:1. This engine is restricted to operate on high alcohol blends

because of its high octane requirement.

Emission results from the 1982 Cressida show the same blend-leaning effect as

the original Cressida. Studies are underway with the intent of defining

modifications that will enhance catalyst efficiency on high percentages of

alcohol.

The high-compression Cressida offers a significant increase in power and performance as wel l as an increase in energy economy of about 10% over that of the

low-compression Cressida, based on the Federal Test Procedures. However, the

high compression Cressida suffers from blend-leaning effects from increased

alcohol blends to a much greater degree than the low-compression engine. This

IV.24
can be seen in Figs. IV.4, IV.5, IV.6. Figure IV.4 presents the CO emissions

versus the methanol blend percentage with the spark held constant at 4 degrees

before top dead center (BTDC). The dotted lines represent the testing performed

with the high-compression engine and the solid lines represent the previous

tests performed with the low-compression engine. The lean shift is apparent in

the decreasing CO emissions as the percentage of methanol is increased. Note

the more pronounced effect in the high-compression engine for the cold and hot

transient segments of the FTP driving cycle. However, the stabilized CO

emissions remained very low for both engines.

The UBF emissions shown in Fig. IV.5 als o decrease with increasing methanol

percentage, again consistent with a leaning of the air/fuel ratio. This effect

is also more pronounced in the high-compression engine.

Figure IV.6 reflects this lean shift in terms of the NOX emissions for each

phase of the FTP. The spark was held at 4° BTDC. The amount in grams per mile

of NOX emissions increases as the percentage of methanol in the fuel increases.

This effect is much more pronounced in the high-compression engine than the lowcompression engine. Also, for each driving phase, the NOX level in the exhaust

was higher for the high-compression engine at equivalent methanol blends,

consistent with other evidence on high-compression ratio engines.

Figure IV.7 shows the effect of retarding the spark on the FTP combined cycle

NOX emissions. Consistent with literature expectation, see Chapter II, the

high-compression engine shows a definite decrease in NOX as the spark is

retarded from 4° BTDC. This effect was also noticeable although to a lesser

degree in the low-compression Cressida. Spark retard had little or no

consistent effect on the UBF or CO emissions nor the fuel economy and driveability.

IV.5g Operating Experience

The 1982 Cressida has now been operated on methanol/gasoline blends for approximately 10,000 miles. One fuel injector has failed. The cylinder head had to be

replaced due to a damaged spark plug. No other problems have surfaced. Driveability and performance are excellent with all methanol blends when powered by

IV.25
.8-r

.6--

ui

.4--

o

o

.2--

SPARK 4 BTDC

DOTTED LINESi High CR

SOLID LINESi Low CR

COLD TRANSIENTrlB A

HOT TRANSIENT •

STABILIZED Q_

+

20 40 60

METHANOL BLEND

9

i

i

100

FIGURE IV. 4 FTP CO EMISSIONS v«. METHANOL BLEND

i
. 35 -r

. 30 --

25 --

LlJ

. 20 4-

CD

\x

u

10 --

. 05 --

. 00

SPARK 4 BTDC

DOTTED LINESi

SOLID LINESs

High CR

Low CR

COLD TRANSIENT-rlB

HOT TRANSIENT

STABILIZED O

4-

0 20 40 60

METHANOL BLEND

80 100

FIGURE IV. 5 FTP HC EMISSIONS v, METHANOL BLEND
4. 0-r

ui 3.0--

\

ow

X

o

z

2.0--

1.0--

0.0

SPARK 4 BTDC

DOTTED LINESi High CR

SOLID LINESs Low CR

COLD TRANSIENT A

HOT TRANSIENT •

STABILIZED O

5 20

\

40

r~

60 80

^

100

METHANOL BLEND %

FIGURE IV. 6 FTP NOx EMISSIONS v«. METHANOL BLEND
1. 6 T

X

O

•z.

1. 4 --

1. 2 --

1. 0 +

0. 8 --

0. 6 --

0. 4 --

0. 2

0. 0

COMBINED CYCLE EMISSION

DOTTED LINESi High CR

SOLID LINESi Low CR

8 BTDC A

4 BTDC A

2 BTDC •

d TDC O

0

FIGURE IV. 7

40 60 80

METHANOL BLEND %

100

FTP NOx EMISSIONS v«. METHANOL

BLEND WITH SPARK RETARD
the standard compression ratio engine and are outstanding for the limited blend

range suitable for the high compression engine.

IV.5h Conclusions

The evidence accumulated with the 1979 and 1982 Cressidas clearly shows that a

current, state-of-the-art system is aval 1 able that w111 al 1ow a broad range

usage in a vehicle without significant adjustments, while maintaining fuel

efficiency, driveability and emissions standards. Admittedly a vehicle will

have somewhat improved performance when set up for a single blend. However, in

times of fuel uncertainty, this project has clearly demonstrated the broad range

blend versatility of the EFI feed-back controlled system and it is recommended

that automotive manufactures note the evidence and incorporate it in their

planning to maximize the worth of their new products to the consumer.

IV. 6 REFERENCES

IV.1 J.F. Nebolon, R. O'Meara, R.K. Pefley, "Pinto Fuel System Conversion To

Pure Methanol And Denatured Ethanol", Final Report, NASA/Grant No. NAG 3-

143, February 1982.

IV.2 J.F. Nebolon, R. O'Meara, R.K. Pefley, R. DiVacky, "The Develoment and

Application Of Kits To Convert A Fleet To Neat Alcohol", V International

Alcohol Fuels Symposium, Aukland, NZ, May 1982.

IV.3 Post Office Alcohol Fueled Vehicles—Final Report in preparation by the

Post Office Department.

IV.4 W.H. Baisley and C.F. Edwards, "Emission and Wear Characteristics of an

Alcohol Fueled Fleet Using Feedback Carburetion and Three-Way Catalyst

System", IV International Alcohol Fuels Symposium, Guaruja, BR,

October 1980.

IV.5 B. Busenthur, "Effects'of Exhaust Gas Recirculation on the Specific Fuel

Consumption and the Emissions of NO, HC, CO, and Aldehyde from a Methanol

IV.30
Fueled, Turbocharged Spark Ignition Engine", Santa Clara University

Mechanical Engineering Research Report, June 1981.

IV.6 S.A. Espinola, A.T. Tamura, J.F. Nebolon, R.K. Pefley, "The Effects of

Fuel Additives on Alcohol Exhaust and Evaporative Emissions", 75th Annual

Meeting of the Air Pollution Control Association, New Orlean, LA, June

1982.

IV.7 U.S. Environmental Protection Agency, "Compilation of Air Pollution

Emission Factors", Supplement No. 5, Appendix D. AP-42,Research Triangle

Park, NC, December 1975.

IV.8 T. Suga, S. Espinola, R. Pefley, F. Cox, J. Al 1 sup, "Performance and

Detailed Emissions of an Electronic Fuel Injection System Which Operates

on any Alcohol/Gasoline Blend", V International Alcohol Fuels Symposium,

Auckland, NZ, May 1982.

IV.9 Alcohol Fleet I Test Program Final Report, California Energy Commission,

August 1982.

IV.10 California Alcohol Fleet II Test Program Final Report, California Energy

Commission, Unpublished Report.

IV.11. S.A. Espinola, R.K. Pefley, "Alternate Fuel Influences on Emissions Test

Procedures," Symposium on Oxygenates in Fuels, American Chemical Society,

Kansas City, MO, September 1982.

IV.12. R. Bechtold, J.B. Pullman, "Driving Cycle Economy, Emissions, and Photochemical Reactivity Using Alcohol Fuels and Gasoline", SAE Paper No.

800260.

IV.13 R.K. Pefley, H.A. Adelman and T.P. Suga, "Methanol/Ethanol/ Gasoline

Blend Fuels Demonstration with Stratified Charge Engine Vehicles,"

California Energy Commission, Contract No. 500-115, Final Report, March

1980.

IV.31
IV.14 W. Balsley and C.F. Edwards, "Emission and Wear Characteristics of an

Alcohol Fueled Fleet Using Feedback Carburetion and Three-Way Catalyst,"

presented at the Fourth International Symposium on Alcohol Fuels

Technology, Guaruja, Brazil, October 1980.

IV.15 R.K. Pefley, J.B. Pullman, T.P. Suga and S. Espinola, "A Feedback

Controlled Fuel Injection System Can Accommodate Any Al cohol -Gasol ine

Blend," presented at the Fourth International Symposium on Alcohol Fuels

Technology, Guaruja, Brazil, October 1980.

IV. 16 R.K. Pefley, et al., "Research and Development of Neat Alcohol Fuel Usage

in Automobiles," Six-Month Progress Report, NASA-Lewis Contract NAG 3-

143, July 1982.

IV.17 T.P. Suga, "Thesis," Master of Science degree in Mechanical Engineering,

University of Santa Clara, to be published.

IV.18 Alcohol Energy Systems, "California Alcohol Fleet Two Test Program,"

Final Report, California Energy Commission Contract No.500-80-054, 1983.

IV.19 R.K. Pefley, T.P. Suga, S.A. Espinola, "Performance and Detailed

Emissions of an Electronic Fuel Injection System which Operates on any

Alcohol/Gasoline Blend," Fifth International Symposium on Alcohol Fuels

Technology, New Zealand, 1982.

IV.32
CHAPTER V

CORROSION AND HEAR MECHANISMS

Based on our earlier work, It was thought that wear would not be a problem with

alcohol fuels using typical gasoline matched lubrication oils. A 1972 Plymouth

Valiant (225 CID), and a 1970 AMC Gremlin, had both been driven 50,000 miles on

neat methanol in typical to severe driving conditions. Oil 'Change intervals

were 4000 miles for both vehicles. In 1979, after 50,000 miles, both engines

were disassembled, and neither engine showed signs of abnormal wear [V.I].

The first evidence of wear problems related to the Santa Clara/DOE program began

with a fleet of five 1979 Ford Pintos operating on neat alcohols. After 10,000

miles, crankcase blowby and oil consumption dramatically increased for three of

these vehicles on neat methanol. Teardown of all three engines revealed

extremely high upper cylinder wear [V.2]. A fleet of 1980 Ford Pintos on

alcohol (known as Fleet I and operated by the state of California) also suffered

from significant upper cylinder wear. The low compression engines were more

inclined to failure than the high compression engines, and those fueled on

methanol showed greater wear rates than those on ethanol. Again, some engines

failed at approximately 10,000 miles [V.3].

Wear studies with dynamometer mounted engines reported in 1977 by Southwest

Research Institute (SwRI) also resulted in high cylinder bore and ring wear when

using methanol [V.4]. Further studies showed that wear rates with methanol were

equivalent to gasoline if the oil sump temperature was above a temperature of

approximately 70 °C. Below 70 °C, the wear rate with methanol , based on PPM

iron in the oil, greatly increased. Wear rates with an oil sump temperature of

46 °C were approximately 10 times the wear rates on gasoline, under equivalent

test conditions [V.5], Furthermore, the addition of 11% water to the methanol

aggravated the wear rate by 3 times compared to anhydrous methanol [V.6].

The development of oil additives specifically for methanol have reduced the wear

rate to "only" twice that of gasoline, based on the standard V-D test [V.7],

This appears to be an improvement, but to effectively deal with the wear

problem, an understanding of the mechanisms which are causing the increased wear

is necessary. Several wel 1-designed experiments at SwRI showed that the wear

V.I
was not caused by oil degradation or excessive oil dilution with water [V.6,8].

As acids are known products of combustion, it was hypothesized that the

increased wear is caused by acid corrosion of the cylinder by the combustion

condensate [V.9], The pH of the condensate would provide a clue, and early

evidence agreed with others [V.10]. A subcontract was signed with Pinnacle

Research Institute, specialists in electrochemistry, to provide analysis and

conduct experiments to investigate the hypothesis [V.ll]. Samples of condensed

pre-catalyst exhaust were obtained from a Pinto on methanol, ethanol and

gasoline. Samples were also obtained from a 1981 Volkswagon Rabbit and a 1981

Ford Escort on methanol with 5.5% iso-pentane, and a 1982 Toyota Cressida on

neat methanol. The pH of the Pinto exhaust condensate on methanol was very low:

1.79 to 1.99. Furthermore, the pH increased with ethanol use and was still

higher for gasoline. Comparison of the Pinto on methanol with the other three

vehicles was also interesting: the pH was higher for the Rabbit, Escort and

Cressida. These pH trends conform to the wear patterns observed with these

vehicles — lower pH values (more acidic) corresponding to higher wear rates. The

condensate was further characterized by titratlon. The results of these tests,

given in Table V.I,

TABLE VI: pH AND ACID-BASE TITRATIONS OF PRE-CATALYST EXHAUST CONDENSATE

Vehicle Fuel pHl Total Acid3

 [HCOOH](N) No of

Pi nto

Pinto

Pinto

Rabbit

Rabbit

Escort

Cressida

MeOH

EtOH

Gas.

MeOH/iso-P

EtOH

Meoh/iso-P

MeOH

1.92

2.24

2.78

2.96

3.44

3.07

3.47

.012

.0058

.0017

.0012

.00041

—

__

Content

.018

.0088

.0063

.0023

—

—

__

.00484

.00114

tests

5

1

2

3

2

2

1

1

 Measured by pH probe

2

 Calculated from pH value

3

 By titration

4

 Assumes HN03 and HCOOH are the only acids present

V.2
show that the activity of hydronium ions, based on the pH reading, is quite

high. The total acid number, from the titrations, indicates the presence of

strong acid(s), as well as weak acid(s). It was hypothesized that the strong

acid is nitric (HN03). and an increased wear mechanism theory was proposed,

based on the following:

1. Methanol produces twice as much water when combusted compared to gasoline.

This causes a higher dewpoint temperature within the cylinder when operating

on methanol.

2. Methanol has seven times the evaporative cooling effect of gasoline, which

increases the time for the cylinder temperatures to rise above the dewpoint after cold starting.

3. Unburned methanol will enter the condensate, while unburned gasoline will

enter the oil. The methanol in the condensate will decrease the surface

tension of the condensate.

4. The condensate may be more acidic compared to gasoline.

Research conducted by Pinnacle Research Institute consisted of four major areas:

1. Definition of the chemical mechanisms which are implied by this wear theory,

and measurement of the following parameters,

2. The corrosion potential of the condensate,

3. The corrosion rate of the condensate,

4. The diffusion rate of the condensate through the lube oil.

V.I A PROPOSED WEAR MECHANISM

The mechanism of wear in methanol fueled engines is based on corrosion due

primarily to hydronium ions (H30

+

) attacking the cylinder wall. These ions will

be present regardless of the type of acids that exist in the condensate. The

most widely accepted mechanism of dissolution of iron in acidic solutions was

proposed by Huesler [V.12], i.e.:

V.3
Fe + OH- < — > Fe(OH)ads + e' (V.I)

Fe(OH)ads + ^ <— > Fe(FeOH)ads (V.2)

Fe(FeOH)ads + OH~ < — >

Fe(OH)ads + FeOH+ + 2e~ (V.3)

FeOH+ + H+ <— > Fe2+ + H20 (V.4)

In this mechanism, reactions 1,2 and 4 are rapid, whil e 3 1s quite si ow and 1s

considered the rate determining reaction. Fe(OH)ads accelerates the dissolution

of iron, and 1s assumed to be an absorbed surface species. Reaction 4 is

strongly pH dependent; if there are high concentrations of hydronium ions HjO"1",

or simply H+, the reaction is fast. However, when the H+ is depleted at the

metal /condensate interface, pH increases and reaction 4 may change course to any

of the fol lowing:

s^ FeOH+ + H20 <—> Fe(OH)2 + H+ (V.5)

FeOH+ + OH" <—> FeO + H20 (V.6)

FeOH+ + OH" <—> Fe(OH)2 (V.7)

The formation of ferrous oxide or ferrous hydroxide would shield the iron from

further attack. However, these products of corrosion, whether FeO, Fe(OH)2,

fe^O^*

 or some other

 compound, are less dense than iron, and would be scraped

away by the piston rings. What is important here is this: if the pH of the

condensate 1s low (£4.00), corrosion will occur.

Since methanol 1s a pure compound, and it or its products of combustion do not

readily react to form any strong acids, some other source of strong acid must be

suspected. The most likely source is N02, which can be converted to nitric acid

by dissolving in water and the following reaction taking place:

3N02 + H20 —> 2HN03 + NO (V.8)

V.4
When the oil sump, and therefore cylinder wall, is cold and below the dewpoint

temperature, water will condense on the oil film. The condensate is a site for

the nitric acid to be formed. There will be two periods when condensation will

occur: when the vehicle is started cold until the lube oil warms up, and after

the vehicle has stopped and the gases trapped in the cylinder cool and condense.

To summarize, nitric acid was proposed as the main source of hydronium ions for

the following reasons:

1. The pH measurements and titrations definitely indicate the presence of a

strong acid, which can not be attributed to a methanol derived acid. Formic

acid would have to be in concentrations approximately 1000 times that of

nitric acid to produce the same pH value.

2. Methanol exhaust has a higher N02/NO ratio and more water than gasoline

exhaust.

3. N02 concentrations are higher in the quench zone, where condensation and

acid formation will occur.

4. Engine wear was observed to be worse at low temperatures, was aggravated by

the presence of water in the fuel, and appeared to be caused by corrosion

[V.8,10,13].

V.2 MEASUREMENT OF THE CORROSION POTENTIAL AND CORROSION RATE

The condensate was characterized by determining its corrosion potential and

corrosion rate using standard experimental techniques. These experiments require

approximately 500 ml of condensate, which would have been impractical to obtain

using the exhaust sampling technique. Therefore, the electrolytic solutions

used in these experiments were surrogates of the exhaust condensate. These

solutions were formulated to have the same pH value as the measured values of

the condensate, and it was assumed that the acids in the condensate were

primarily nitric and formic (HCOOH). The composition of the electrolytic

solutions is given in Table V.2.

V.5
TABLE V.2: COMPOSITION OF THE ELECTROLYTIC SOLUTIONS

Electrolyte [HN03](N) [HCOOH](N) pH

Number1

1

2

3

4

52

6

3

7

4

8

9

10

11

,122

133

0.0200

0.0180

0.0160

0.0140

0.0160

0.0160

0.0160

0.0027

0.0020

0.0013

0.0006

0.0013

0.0013

0.0013

0.000

0.002

0.004

0.006

0.004

0.004

0.004

0.000

0.0007

0.0014

0.0021

0.0014

0.0014

0.0014

1.70

1.74

1.78

1.83

1.77

1.75

1.73

2.53

2.68

2.80

2.95

2.78

2.74

2.73

1 Electrolytes 1-7 were tested with electrode of

Pinto block material. Electrolytes 8-14 were

tested with electrode of Rabbit block material.

2 Contains three volume percent methanol

3 Contains six volume percent methanol

4 Contains ten volume percent methanol

A standard three-electrode electrochemical cell was used for the experiments.

The working electrode was a cast iron rod 0.2" in diameter. The rods were

machined from the engine blocks of a Ford Pinto and VW Rabbit, and were encased

in an epoxy sheath so that only one end was exposed to the electrolyte. The

other end of the rod was attached to a rotating disc el ectrode(rde) shaft. A

silver/silver chloride (Ag/AgCl) half-cell was used as the reference electrode.

V.6
A platinum gauze of 6.2 in2 exposed surface area was used as the counter

electrode. The electrochemical cell Is illustrated in Fig. 1 [V.ll].

The corrosion potential or rest potential , EC, was determined as the voltage

produced at the elect rode (engine block)/electrolyte interface. It is a measure

of the tendency of the metal to corrode and was found to be a linear function of

the pH of the electrolyte over the range of pH tested. The results from the

experiments are shown in Fig. V.2. The change in Ec with respect

to pH is a constant, and can be expressed as:

dEc

------- = _60 mv/decade

dpH

In later experiments, a potential was applied across the cell and the current

was measured. The total potential equals the rest potential, Ec, plus the

applied, or over potential, . Jhe anodic polarization curves are plots of the

logarithm of current versus overpotential. Figure V.3 is a typical plot, which

is based on the tests with electrolyte 1.

The linear portion of the curves are dominated by the kinetics of the metal

dissol ution and are often referred to as Tafel lines. The Tafel lines can be

expressed as:

i » ic[ exp(2.37/b)]

where: ic = corrosion current density obtained by linearly

extrapolating to intercept Ec(rj = o)

and b = Tafel slope = d ft

dlog i

The Tafel slope and pH dependence of the corrosion potential are consistent with

studies of the corrosion of iron in acid media [V.14,15].

V.7
Ag/AgCl

REFERENCE

ELECTRODE

Ag WIRE

ROTATING DISK

ELECTRODE

~-WORK ING

ELECTRODE

—Pt COUNTER

ELECTRODE

FIGUREV1: THE ELECTROCHEMICAL CELL
-O. 6

J

Iz

u

I-o

Q_

z

O

M

tn

o

o:

o:

o

u

-o. y_.

-o. s_.

-o. 9

_.x PINTO BLOCK ELECTRODE

o RABBIT BLOCK ELECTRODE

O 1. O 2. O

PH

3. O 4. O

FIGUREV.2 : CORROSION POTENTIAL vs. PH
1. O

IZ

liJ

tt.

CL

D

U

10 ..

01

--/ o

-- o

-- o

24 48 72 96

OVERPOTENTIAL CmV>

120

FIGUREV3: TYPICAL ANODIC POLARIZATION CURVE
The corrosion rate based on 1C can be calculated using the following equation:

1py = 1C t MW (1)

n

 F (2.54)

where: ipy = inches penetration per year (1n/yr)

ic = corrosion current density (A/crri2)

t = number of seconds per year (3.1536 x 10; sec/yr)

MW = molecular weight of the corroding metal (55.847 gm/mole for

i ron )

n = number of charges transferred, or the oxidation state of the

respective corrosion products (n = 3 for iron)

F = Faraday constant (96,500 A sec/equiv)

f = density of the metal (7.89 gm/cm3 for iron)

The corrosion rates based on these tests were calculated to range from 0.009 to

0.059 ipy. Of course, these rates are an overestimate, because in actual

conditions the condensate exposed to the cyl inder wal 1 is much less, and the

acid would be depleted as corrosion proceeds. Nevertheless, these calculations

are useful in comparisons among different condensate compositions.

V.3 MEASUREMENT OF THE CONDENSATE DIFFUSION RATE

The condensate diffusion rate was measured by immersing an oil coated carbon

steel electrode into an electrolytic solution. When the condensate penetrated

the oil film and reached the electrode, the open circuit voltage across the

electrode became the rest potential. The elapsed time from the immersion of the

electrode into the condensate until the detection of Ec is called the

transition time, measured in seconds.

The oil film thickness was varied by dipping the electrode into an oil (20W-40

weight SF-CC oil, by Conoco), and rotating the electrode at a given speed for at

least 30 minutes. The thickness was determined by measuring the weight increase

of the electrode. A plot of the data indicates that oil film thickness is

inversely proportional to rotating speed (Fig. V.4).

V.ll
2,0

•\

CO

CO

CD

c

JC

•M

_£=

•I—I

M—

«O

1,0

0

0,5 1,0 1.5 2,0

1/w x 10"3 (rotation speed.,

Figure V,4: Correlation of Oil Coating Film Thickness

with Rotation Speed
The electrolyte was composed of 0.016 M^ HNOa and 0.004 M_ HCOOH which was modified to produce four test electrolytes by adding 0, 3, 6, and 10 volume percent

methanol. The results of these tests, transition time for four different

solutions and three different oil film thicknesses, a/e given in F1g. V.5.

These results are truly significant. Methanol in the condensate can greatly

increase the ability of the condensate to penetrate the oil film on the cylinder

wall and initiate corrosion. Of course, unlike the experimental conditions the

oil film in the cylinder is not static; the film is wiped and replenished with

each passing of the piston rings. However, with the engine running more

methanol laced condensate will reach the cylinder, continuing the corrosion

until the cylinder wall rises above the dewpoint temperature.

V.4 FURTHER ANALYSIS OF THE CONDENSATE

In parallel with the corrosion tests, more exhaust condensate samples were

obtained and analyzed for pH. Ten different vehicle/fuel combinations were

tested: Ford Pintos, on methanol, ethanol (CDA-20), and gasoline, VW Rabbits,

on ethanol and 94.5% methanol/5.5% iso-pentane, a Ford Escort, on the same

methanol/iso-pentane blend, and a Toyota Cressida on M80, M90, and M100 (M100

with a high and low CR engine). These samples were obtained from pre- and postcatalyst exhaust and blow-by gases directly from the crankcase, all while the

vehicle engines were idling. This data was taken from in-use vehicles over a

period of a year, and all the variables were not well controlled. However,

several trends can be observed in Fig. V.6.

Post catalyst exhaust was generally at least one pH number higher compared to

pre-catalyst exhaust. Blow-by exhaust condensate from the Escort had the same pH

as the pre-catalyst sample, while the Pinto blow-by showed a definite increase

of about 1.3 above the pre-catalyst exhaust. All the vehicle/fuel combinations

except for two showed a consistent value of pH. Note that methanol condensate pH

values were on both sides of the gasoline data. What besides fuel affects pH

values is not completely understood at this point. However, one possible answer

was revealed in recent work.

V.13
90

SO ..

70 ._

U

UJ

8 60

UJ

I 50

fo '

(/) 3O

2

I- 20 ..

1O ..

O%

3%

1.5 2. O 2. 5 3. O

OIL FILM THICKNESS <CM X

3. 5

FIGUREV.5 TRANSITION TIME VS. OIL

FILM THICKNESS FOR -COMPENSATE.WITH

VARIOUS METHANOL PERCENTAGES

I
198O FORD PINTO

METHANOL

19BO FORD PINTO

ETHANOL

1978 FORD PINTO

, . GASOLINE

1981 FORD ESCORT

METHANOL/ISO-P

1981 VW RABBIT

METHANOL/ISO-P

1981 VW RABBIT

ETHANOL

1982 TOYOTA CRESSIDA

LCR M100

1982 TOYOTA CRESSIDA

HCR M1OO

1982 TOYOTA CRESSIDA

HCR M9O

1982 TOYOTA CRESSIDA

HCR M8O

PRE-CATALYST

POST-CATALYST

CRANKCASE BLOWBY

If

4-

1. 0 2. 0 3. 0

pH

4. 0

. 13

f * f

6. 20

I

5. 6O

5. 0

FIGUREVGs EXHAUST CONDENSATE PH VALUES

FOR VARIOUS VEHICLE/FUEL CONFIGURATIONS
Pre-catalyst samples were obtained and analyzed for pH and the presence of ions

by ion chromatography. The ions which were detected were nitrate, nitrite,

formate, .sulfate, and; chloride, which were assumed to exist as nitric, nitrous,

formic, sulfuric, and hydrochloric acid, respectively. The concentrations of

these ions are given in Table V.3.

TABLE V3: ION CONCENTRATIONS(mg/L) AND pH VALUES

OF PRE-CATALYST EXHAUST CONDENSATE

Vehicle: 1980 Ford Pinto(49 state)

Fuel: Methanol Ethanol(CDA-20)

- Sample 1 Sample 2 Sample 1 Sample 2

Formate 21. 1.2 62. 41.

Nitrate(as N) 15. 6.6 9.4 11.

Nitrite(a s N) 0.56 0.23 0.66 3.1

Sulfate 8.6 24. 9.3 15.

Chloride 9.4 6.5 9.9 3.7

pH 3.49 4.15 4.12 4.35

Nitrate was a major part of the ions detected; however, there were equally

large amounts of sulfate and chloride ions. This data indicates fuel contaminants, specifically sulfur and chlorine compounds, as a source of corrosive

agents, at least in these tests. Other groups have reported instances of fuel

contaminants which c-ould affect wear rates [V.7,10,16], Unfortunately, analysis

of the fuel was not a part of this study, but one conclusion has been reached:

analysis of the fuel prior to use is necessary to maintain quality control and

to allow accurate interpretation of wear data.

V.5 CONCLUSIONS AND FUTURE WORK

Follow-on studies will focus on three aspects of the alcohol fueled engine

corrosion and wear problem.

V.16
1. Future exhaust condensate experiments will concentrate on improving the

procedure for sampling and analyzing the combustion condensate toward better

correlation of pH and ion evidence with vehicle operating conditions. Fuel

quality will be part of the data and the in-cylinder condensate will be

compared with exhaust and blowby condensate.

2. Existing computer combustion reaction kinetics modeling will be used to

interpret the condensate evidence and provide clues to suppression of the

acidic nature of the condensate.

3. The accelerated corrosion measurement procedures will be modified as

necessary based on the new condensate evidence and restudied to obtain clues

as to how best block the corrosion attack if the corrosive elements cannot

be suppressed.

The dominant mechanism proposed for corrosion and wear of methanol fueled

engines is based on the hydronium ion, which is provided by acids in general and

involves the following features of alcohol combustion:

1. There is more water formed by combustion and more time for it to condense

relative to gasoline because of the high dew point temperature of the

exhaust gases and high heats of vaporization of the alcohols.

2. With pH values of less than four having been measured for blowby and precatalyst exhaust condensate, electrochemical experiments show that corrosion

of cast iron will occur when exposed to these solutions.

3. The diffusion of the condensate through the oil film is much faster when

.small percentages of methanol are present in the condensate in comparison

with condensate from gasoline engines. .

4. Based on all the evidence to date, an acidic corrosion wear mechanism

involving the hydronium ion is most likely, particularly when considering

the conditions for highest wear rates. Whether the dominant acid is nitric

or other remains to be proven but some of the most severe corrosion cases

are clearly linked to strong acids by the condensate evidence.

V.17
V.6 REFERENCES

1. R. K. Pefley; 1981; "Characterization and Research Investigation of Alcohol

Fuels in Automobile Engines;" DOE Contract No. DE-AC03-78CS; Final Report;

pp 121-124.

2. W.H. Baisley, C.F. Edwards; " Wear Characteristics of Fleet Vehicles Operating on Methyl Alcohol;" SAE Paper No. 811202.

3. R. Ernst, R.K. Pefley, F.J. Weins ; 1983; "Methanol Engine Durability;" SAE

Paper No. 831704.

4. E. C. Owens; 1977; "Methanol-Fuel Effects on Spark Ignition Engine Lubrication and Wear; " II International Symposium on Alcohol Fuel, Technology

Methanol and Ethanol; 2-6.

5. E.C. Owens, H. Marbach, Jr., E. Frame and T. Ryan; 1980; "Lubrication

Requirements for Alcohol-Fueled Spark Ignition Engines;" IV International

Symposium on Alcohol Fuel Technology; Vol. 2 (B-66); pg 691-700.

6. E.C. Owens, D.W. Naegeli, H.W. Marbach, T.W. Ryan III and E.A. Frame;

1983;"Approaches to Controlling Wear in Alcohol Fueled Engines;" V International Alcohol Technology Symposium; Vol. 2 (C3-5);pg 443-450.

7. H.W. Marbach,Jr., E.C. Owens and E.A. Frame; 1983; "Technical Report for

Resolving Wear Problems in Alcohol-Fuel ed Engines;" Technical Report No.

SwRI-7152/1; pp 20.

8. T.W. Ryan III, D. Naegeli, E.C.Owens, J. Barbee; 1981; "The Mechanisms

Leading to Increased Cylinder Bore and Ring Wear in Methanol-Fueled SI

Engines;" SAE Paper No. 811200.

9. R.K. Pefley; "Alcohol Fuel Corrosion and Wear Effects;" 1982; presented at

the United Nations Industrial Development Conference; Dehra Dun, India.

V.18

I
10. H.C. Wolff , D. Smith, P. Fry; 1982; "Motor 011 Development for Methanol

Engines;" V International Alcohol Fuels Technology Symposium; Vol 2(C3-8);

pg 467-474.

11. Pinnacle Research Institute,Inc.; 1983; "The Corrosion of Metals in

Methanol-Fueled SI Engines;" Final Report, 1n preparation.

12. K.E. Huesler; 1958; "Z. El ectrochem;" Vol 2(12); pg 1371.

13. D.H.T. Milliard, B.J. Hewlett, S. Chaibongsai; 1982; "Crankcase Oil Development for Alcohol Fueled Engines; "V International Symposium Alcohol Fuel

Technology; Vol 2(C3-4);pg 435-442.

14'. 6.M. Florianovich, L. Sokolova, Y.M. Kolotyrkin; 1967; "Electrokhemya;" Vol

3(9); pg 1027.

15. T.D. tytner, A.L. Rotinyan; 1966; "Electrokhemya;" Vol 2(12); pg 1371.

16. H.W. Marbach, Jr., E.A. Frame, E.G. Owens, D.W. Naegeli; B.D. Wielgos ; 1983;

"The Effects of Lubricant Composition on SI Engine Wear With Alcohol Fuels;"

SAE Paper No. 831702.

V.19
CHAPTER VI.

PHOTOCHEMICAL SNOG CHAMBER STUDIES

The 100 cubic foot photochemical smog chamber was transferred to the University

of Santa Cl ara from Bart!esvi 1 1 e Energy Technology Center (BETC) i n January

1977. During the first year it was installed and supporting equipment and

instrumentation were acquired. Continuous measurement techniques were developed

for total hydrocarbons, oxides of nitrogen, ozone and formaldehyde. The

original aluminum- and Pyrex- walled chamber used 84 blacklights and 18 sunlamps

for solar simulation.

The smog chamber and its supporting equipment were unused from 1980 through mid1982 when renewed interest in methanol's photochemistry emerged and an improved

program and experimental plan were devised. The rationale for conducting

additional photochemical smog chamber studies of urban hydrocarbon and methanol

exhaust surrogates is to provide validation for their chemical kinetic

mechanism. The chemical kinetic mechanism is subsequently modeled along with

the sunlight, temperature, physical transport, emissions mixing and dilution

processes which occur in a vertical column of air as it traverses an urban path.

This Empirical Kinetic Modeling Approach (EKMA) is endorsed by the U.S.

Environmental Protection Agency for use in the formulation of 1982 and

subsequent air quality State Implementation Plans. Therefore, the methodology

seemed most appropriate for an assessment of methanol fuel's impact in the Los

Angeles Air Basin. However, because this work is independently funded by a

multiclient group and is still underway, the present report concerns the

restoration efforts, characterization studies and experimental progress achieved

with the photochemical smog chamber apart from the EKMA modeling study.

VI.1 FACILITY UPGRADE

The restoration of the photochemical smog chamber facilities required several

months of effort. The chamber was first relocated adjacent to the laboratory's

electrical power bus. Electrical power connections and circuit breakers were

needed to supply the 1.5 kilowatts consumed by the 84 blacklights and 18

sunlamps. Additional power for the rack-mounted instrumentation package was

VI.1
also supplied. The electric furnace which contains a catalytic combustor for

air purification also required an independent power supply.

The clean air supply for calibration zeroes and chamber dilution purposes was

redesigned incorporating a new blower. An ozone generating ultraviolet lamp was

.purchased and installed downstream from the blower. This was intended to

oxidize any nitric oxide into nitrogen dioxide. The ozonated air supply was

next routed to a dust filter and then into a potassium permangenate (Purofil )

adsorbent canister. The removal of the majority of inorganic material was thus

accomplished. The clean air supply was next routed to the catalytic combustor

for removal of organic contaminants. The furnace temperature and air flowrate

parameters were adjusted to produce a clean air supply of about ten liters per

minute. A portion of that supply was pumped into a combination instrument

calibrator and chamber dilution rate controller. The remaining supply of clean

air can be vented into the laboratory; or it can be diverted into the chamber

for rapid dilution and flushing of old experiments.

VI.2 INSTRUMENTATION

The instrumentation set which was previously acquired with DOE support, was

removed from storage and rack mounted. The instruments consist of:

(1) Monitor Labs Model 8500 calibration controller

(2) Beckman Model 400 Total Hydrocarbon (THC) Analyzer

(3) Dasibi Model 1003 AH ozone monitor

(4) Monitor Labs Model 8440 Nitrogen Oxides (NOx) Analyzer

A dual pen chart recorder and two data multiplexers were also installed to

record the raw data from the experiments. Separately located was a CEA 555

automated wet chemistry formaldehyde analyzer and its companion calibration

standard. A formaldehyde permeation device was located within a glass U-tube

which was immersed in a 100°C temperature controlled silicone ,oil bath.

A Hewlett-Packard Model 5721 gas chromatograph was located adjacent to the smog

chamber. Air samples were pumped from the smog chamber and. into a glass-woolfilled sample trap. The 1/8 inch stainless steel trap was submerged in a liquid

VI. 2
argon bath in order to adsorb and collect an increased concentration of

hydrocarbons for measurement. A six port valve was usekl to switch the sample

loop into the nitrogen carrier gas stream. The cryogenic trap was immediately

piunged into a boi1 ing water bath to rapidly desorb the col lected sample. A

Hewlett-Packard model 3380 integrator was used to assist in quantifying the

measured gas concentrations. The lower detectable limit for propene was

measured at 2 to 3 parts per billion carbon using the trapping technique

described. The linearity of the propene measurements was checked by a series of

analyses after repeated addition of a fixed volume of propene to the chamber.

Good linearity was found for the sample trapping, desorbtion and gas chromatography procedures used.

VI.3 CHAMBER GAS SAMPLING

The sampling system originally obtained continuous gas samples from several

locations around the chamber. Stainless steel, glass and Teflon were the only

materials in contact with the samples. Some of the instruments chemically

altered the samples so that they had to be discarded after use. However, the

unaltered total hydrocarbon analyzer and the ozone analyzer samples were

recirculated back into the chamber in order to minimize sample 1 oss di1ution

effects.

The preceding is a description of the experimental facilities when the first

series of characterization studies and the multi-client supported methanol

photochemistry study commenced in late 1982 and during the first six months of

1983. The characterization studies and initial methanol photochemistry

validation studies were found to be subject to chamber wall contamination and

aging effects about one month or less after the aluminum and Pyrex surfaces

were cleaned. The magnitude of the contaminant effects was substantial enough

to strongly interfere with the reaction rates of the hydrocarbons and NOx under

study. The severity of the problem was not fully assessed until mid-June of

1983. It was then decided to upgrade the smog chamber walls with the addition

of an inner flouro-ethylene-polymer (FEP) Teflon film bag. In that way the

hydrocarbon and NOx surrogate mixtures would only be in contact with their FEP

film enclosure. The FEP film was purchased and a 100 ft3 bag was constructed to

conform with the aluminum and glass walls of the original chamber. The FEP film

VI.3
bag was supported with detachable Velcro strips along its two upper sides.

The effective volume of the chamber was reduced to 70 ft^ since the two ends

were unsupported and partially collapsed.

The clean air supply was also upgraded with the addition of a high efficiency

particulate air (HEPA) filter located just after the catalytic combustor. A

fraction of the clean filtered air was permanently routed into the buffer-zone

air surrounding the FEP film bag, but within the aluminum and glass chamber

walls. This was a further precaution against the diffusion of hydrocarbon and

NOx contaminants from the external laboratory air into the FEP film bag.

The sampling system was also modified at that time; One inlet and one exit

bulkhead union were fitted through the aluminum chamber wal1 s. A large Teflon

plug was drilled and tapped and located so that the FEP film wall was sandwiched

between the Teflon plug and the aluminum chamber wall for sealing purposes.

The four sample lines were routed to their respective instruments outside of the

chamber. The recirculated sample (ozone and THC) was returned via the bulkhead

union on the opposite end of the chamber. An external one half inch Teflon

tube and mixing pump were also added to help homogenize the contents of the bag

when concentrated hydrocarbon and NOx reagents are added. This completed the

modifications to upgrade the clean air and sampling systems and reduce chamber

wall effects.

VI.4 CALIBRATION STUDY

A repeat series of chamber characterization studies was performed next. These

experiments were modeled by a subcontractor (Systems Applications, Inc. (SAI) of

San Rafael, California) in order to confirm the performance of the chamber

against their knowledge of other chambers and the known chemistry of particular

experiments. The most fundamental chamber characteristic to model is simple

dilution of a hydrocarbon with a measured flowrate of clean air. Given the

chamber volume and the-dilution flowrate one can readily calculate the

theoretical decrease in concentration C(t) as a function of time:

VI.4
C(t) = C(init) * exp (-b)t (VI.1)

dilution flowrate

where b =

chamber volume

t = time elapsed since initial concentration: C(init)

Figure VI.1 shows the agreement between the calculated and measured chamber

di1ution rates.

A second fundamental chamber characterization involved the increase in chamber

temperature after turning on the lights. Interpreting and modeling of the

experimental results is enhanced by a knowledge of the temperature profile. A

typical temperature profile over time is presented in Fig. VI.2. The curve

maxima are due to the evening cooling effect of the surrounding laboratory air.

Some of the chemical kinetic rates are sensitive to temperature effects although

these are generally small in magnitude.

A series of special experiments were designed with guidance from SAI to measure

the reactivity contribution of the chamber's walls. These experiments involved

the use of NOx, CO, and acetaldehyde. The acetaldehyde experiments used only

acetaldehyde near the ppm level to determine the rate of NOx off-gassing from

the chamber walls. Such off-gassing could be important to experiments limited

by NOx concentrations. For example, the ozone peak under high hydrocarbon to

NOx ratio conditions is usually determined by the point when NOx approaches

zero. If nothing is purposely added to a chamber and NOx measurements are made,

very little if any NOx is seen because there is usually some radical concentration available to convert any residual NOx to N02

 and

 then any hydroxyl radicals

(OH) would rapidly convert the N02 to nitric acid. Hence the NOx measurements

would show only a trace steady-state sum of NO and N02 (i.e. NOx) even though

NOx might be off-gassing at a significant rate. Acetaldehyde provides some

radicals by photolysis and OH radicals produce peroxyacetyl radicals from

acetaldehyde. Hence, any NOx off-gassing leads to peroxyacetyl nitrate (PAN)

formation. Most chemiluminescent N0£ measurements include interference by PAN.

Hence any off-gassing of NOx in the presence of acetaldehyde and ultraviolet

VI.5
1200-

TOTAL

PPB'S

CARBON 1000-

DIVIDED

BY 9

VS. egg

CALCULATED

DILUTION

RATE "D"

600 +

PARTS

PER

BILLION

200- -

DATE : AUG 04 TYPE : DILUTION

D MEASURED VALUES

— PREDICTED VALUES

120 240 360

MINUTES OF DILUTION

480 600

FIGURE VI.1 : EXPERIMENTAL VS. CALCULATED SMOG CHAMBER DILUTION COMPARISON
DATE sOCT 20 TYPE s TEMPERATURE

aa

38-- a

TEMPER- 34<

ATURE

DEGREES 32.

CELSIUS

30--

28

aaaaaaaaaaaa a

aa a aaaaaaa a

a a an

nnnnnnnn

n nnnn n n n n n n

- - a nnn

n

•a n

n n n n n m

-- n

a = TEMPERATURE INSIDE BAG

n = SURROUNDING TEMPERATURE

- -n

A I 1 1 1 1 1 1 1 1 1

120 240 360

MINUTES OF IRRADIATION

480 600

FIGURE VI. 2 i TYPICAL TEMPERATURE PROFILE DURING SMOG CHAMBER EXPERIMENT
light would show a buildup of PAN and could be measured with a normal N02

detector. Figure VI.3 shows the acetaldehyde experiments performed recently

which were modeled with acetaldehyde chemistry and a NOx off-gassing rate of

only 0.04 ppb per minute. The modeled rate of NOx off-gassing is a low enough

value to not seriously affect the results of this study and its value is

consistent with other smog chambers. {

Two types of experiments have been used to determine the organic and radical

reactivity of the smog chamber. In the first experiment about 0.5 ppm of NO is

the only precursor added to the chamber. Any observed conversion of NO to N02

would be caused by trace organics reacting in the presence of some source of

radicals. The second experiment is the same except 50-100 ppm carbon monoxide

(CO) is added. The CO supersedes any trace organic reactivity because CO reacts

with OH to generate C02 and the hydroperoxy radical (H02) which in turn reacts

with NO to give N02 and return the OH. Unlike most organics CO does not produce

any products which can photolyze and provide radicals to sustain the smog,

chemistry associated with conversion of NO to N02- Hence the NO to N02.

conversion in the CO experiment amplifies the source of radicals in the chamber.

The experiment without the added CO shows the result of both the chamber

radicals and the organic reactivity.

Figures VI.4 and VI.5 show the NOx and NOx plus CO experiments which were fitted

using an assumed initial background level of formaldehyde of 0.04 ppm plus a

trace of formaldehyde coming off the walls at a rate of 0.0002 ppm per minute,

plus a conversion of N02 to nitrous acid (HONO) with a rate constant of 0.00024

per ppm per minute. This last value corresponds to a recently determined

heterogeneous pathway studied by Sakamaki et al. [VI.1]. Such a reaction

provides radicals from the subsequent fast photolysis of HONO to OH and NO. The

use of assumed trace level s of formaldehyde is based on the work of Whitten,

Kill us, and Hogo [VI.2], Such low levels of organic and radical species required

to fit the chamber characterizations correspond to one of the lowest seen in

smog chambers. Some recent data from a similar chamber at the University of

California at Riverside required similarly low level s to simulate such

experiments. The radical source strength corresponding to the formaldehyde offgassing and heterogeneous N02 to HONO conversions is about 0.05 ppb per minute,

VI .8
DATE : NOV 22 TYPE s NOX OFFGAS

1200-r

1000--

800--

600--I

OZONE

ACETALDEHYDE

NOx

120 240 360 480 600

MINUTES OF IRRADIATION

FIGURE VI. 3 « ACETALDEHYDE CHARACTERIZATION TO MEASURE NOx OFFGASSING RATE
1200-r

NOX

NO

N02

OZONE

THC/3

1000--

800 --

PARTS

PER

BILLION

600--

200--

DATE i JUL 19 TYPE s CLEAN AIR,

4QO--JffiX

" NO.

-- N02

•=THC

OZONE

120 240 360

MINUTES OF IRRADIATION

480 600

FIGURE VI. 4 : CHARACTERIZATION EXPERIMENT USING CLEAN AIR AND NOX
1200-r

NOX

NO

N02

OZONE

1000--

800--

600 -•

DATE tOCT 12

DATE iQCT 13

TYPE i 50 PPM CO

TYPE i 100 PPM CO

480 720

MINUTES OF IRRADIATION

960 1200

FIGURE VI. 5 : CHARACTERIZATION EXPERIMENTS USING CARBON MONOXIDE AND NOX
well below the typical 1 ppb per minute or greater generated from the organics

and their products used in the regular experiments.

A very important characteristic of the smog chamber is its ultraviolet light

intensity. The solar simulation was provided by 84 blacklights (F30T8/350 Bl)

and 18 sunlamps (FS20). The lamps were individually replaced over an 840 hour

life cycle in order to control aging effects. The light must pass through

"pyrex" glass windows which selectively attenuate the shorter wavelengths in

the ultraviolet spectrum. The resulting spectral distribution of light

intensity is analogous to that produced from sunlight after it passes through

the earth's atmosphere.

Two photolysis reactions are of key importance in driving the photochemistry

which produces urban ozone pollution. The first reaction is formaldehyde

photolysis to yield radical products:

hv

HCHO > 2H02 + CO (VI.2)

These products initiate radical chain reactions which further oxidize other

hydrocarbons and eventually convert nitric oxide into nitrogen dioxide.

NO + H02 —-> N02

 + OH

 (VI-

3

)

A series of propene/NOx characterization studies were performed in order to

better understand the formaldehyde photolysis rate constant. Past modeling

experience by SAI has shown this to be the most important factor for achieving a

close simulation of observed chamber data (VI.3). Therefore, the formaldehyde

photolysis constant was determined by fitting the propene experiments: these are

shown in Fig. VI.6. The key formaldehyde constant for photolysis to radicals so

determined was 0.0025 per minute, a fairly high value but not inconsistent with

the mixture of blacklights and sunlamps used in this chamber. The photolysis

constant to stable products from formaldehyde is not important to smog

chemistry, but an equal value of 0.0025 per minute was used to be consistent.

The smog chamber's formaldehyde photolysis rate constant is compared in Fig.

VI.7 with the EPA's computed solar value for Los Angeles on June 26. There is a

diurnal variation in sunlight intensity for formaldehyde photolysis which

contrasts to the higher intensity but relatively constant chamber lights.

VI.12
1200-

NOX

NO

NQ2 1000-

OZONE

PROPENE

800-

DATE*AUG31 TYPE » C3H6/NOX PPBC/NOx i 3.06

PARTS

PER

600

BILLION 40Q--

200--

240 360

MINUTES OF IRRADIATION

480

OZONE

600

FIGURE VI. 6 : PROPENE & NOX CHARACTERIZATION FOR FORMALDEHYDE PHOTOLYSIS
0.30-r

DISSOCIATION 0.25

RATECONSTANT

FOR

 0.20+

HCHO*100

<MIN

-1 0. 15- -

0.10- -

Q. 05- -

CHAMBER PHOTOLYSIS RATE CONSTANT

URBAN PHOTOLYSIS RATE CONSTANT

f f f

f

H 1 1

0700 0900 1100 1300

TIME OF DAY (JULY 26)

1 1 1 1 1 1 1

1500 1700

FIGURE VI. 7 : EXPERIMENTAL & URBAN FORMALDEHYDE PHOTOLYSIS RATE CONSTANTS
The second photolysis characteristic of key Importance 1s the dissociation rate

constant for nitrogen dioxide:

hv

N02 > NO + 0 (VI.4)

The method of Wu and Nikl [VI.3] was used to determine the chamber's effective

N02 dissociation rate. The measured value of 0.25 min-1 is compared in Figure

VI.8 to the EPA's computed solar value for Los Angeles on June 26 [VI.4]. It is

evident that the smog chamber's lights are less intense than mid-day urban

sunlight for the photolysis of N02-

 The sm°9 chamber, however, is used only to

experimentally validate the chemistry of the hydrocarbon, alcohol and NOx surrogates under study. In this study the lower intensity but more constant chamber

lights are treated accordingly when modeled. The chemical kinetic mechanism is

subsequently used in an urban model which accurately represents the diurnal

variations in the two photolysis rate constants discussed above.

The hydrocarbon surrogate mixture used in this study was designed to represent

the current chemical classes of hydrocarbons in the South Coast Air Basin. It

includes eight surrogate species rather than the two or four hydrocarbon

surrogate blends used in previous studies. Table VI.1 lists the organic

compounds and their relative carbon percentages within the surrogate mix.

Table VI.1 : Urban Hydrocarbon & Methanol Exhaust Surrogates

HYDROCARBON SURROGATES

ethene

propene

isobutene

n-butane

n-pentane

2,2,4-trimethyl pentane

toluene

m-xylene

METHANOL SURROGATES

methanol

isobutene

CARBON %

5

5

15

15

20

15

12.5

12.5

90

10

VI.15
0.6-r

DISSOCIATION

RATE 0-5

CONSTANT

FOR

N02 0.4.

-1 0.3--

CMIN )

0.2--

0.1--

n

DIURNAL CHANGE IN RATE CONSTANT

n

n

n

n

CHAMBER PHOTOLYSIS RATE CONSTANT

0700 0900 1100 1300

TIME OF DAY (JULY 26)

1500 1700

FIGURE VI. 8 : EXPERIMENTAL & URBAN NITROGEN DIOXIDE PHOTOLYSIS RATE CONSTANTS
VI.5 EXPERIMENTATION

The experimental design is outlined in Table VI.2. The surrogate mixture of

hydrocarbons was tested at three overall hydrocarbon (PPM carbon) to NOx ratios.

Three sets of experiments were performed at each ratio: one with the full

mixture (baseline), one with only two-thirds of the concentration (blank

substitution), and one at two-thirds plus the remaining carbon concentraton

substituted by a mixture of methanol and isobutene in a nine to one molar ratio

(methanol substitution). Isobutene reacts rapidly to generate formaldehyde so

the addition was equivalent to adding formaldehyde at the ten percent level to

simulate exhaust from methanol vehicles.

Table VI.2 : Experimental Design & Initial Surrogate Concentrations

Experimental

Conditions

Baseline

Blank

Substitution

Methanol

Substitution

PPM Carbon

to Nox Ratio

3

9

27

2

6

18

3

9

27

NOx

1.2

0.4

0.4

1.2

0.4

0.4

1.2

0.4

0.4

8-Surrogate

Urban Blend

3.6

3.6

10.8

2.4

2.4

7.2

2.4

2.4

7.2

Methanol Exhaust

Surrogates

0

0

0

0

0

0

1.2

1.2

3.6

An example of a complete set of experimental measurements is plotted in Fig.

VI.9. The initial reactant concentrations are shown at time zero on the left of

the figure. Each of the eight urban surrogate hydrocarbons is represented by a

letter symbol which is identified in the key at the upper right. Initial oxides

of nitrogen (nitric oxide and nitrogen dioxide) are also plotted along with the

photochemical products formaldehyde and ozone.

The observed change in concentration of each species is the result of both

controlled dilution and photochemical reactions. The purpose of this experiment

is to validate an evolving chemical kinetic mechanism which describes in detail

the chemical reactions and their rates within the system under study. The

chemical kinetic mechanism is then incorporated into a model of the physical

VI.1 7
EXP DATE : JUL 28. 1983

TYPE t 8S BASELINE

PPBC/NOx : 8.86

C4H8

C4H10

ACET

P C5H12

C8H18

120 240 360

MINUTES OF IRRADIATION

480

FIGURE VI. 9 : REACTANT CONCENTRATIONS AND SMOG CHAMBER PRODUCTS OVER TIME
processes and emissions inventories of a specific urban setting for ozone

modeling purposes. If the methanol substitution experiments produce results

similar to the two-thirds level without methanol then one would conclude that

methanol was inert. If the substitution experiments are the same or more

reactive than the full concentration experiments then one should conclude that

the methanol was as reactive or more reactive than the urban mixture.

The comparative set of experimental results at the lowest ppm carbon to NOx

ratio concentration is presented in Fig. VI.10. Only nitric oxide, nitrogen

dioxide and ozone are plotted for these experimental overlay comparisons. The

evidence indicates that the methanol substitution experiment was very similar in

reactivity to the blank substitution case. Therefore, at this low ppm carbon to

NOx ratio the methanol exhaust surrogates were relatively inert and contributed

very little to ozone production.

The nine to one ppm carbon to NOx ratio evidence is presented in Fig. VI. 11.

The methanol substitution experiment was intermediate in reactivity at this

reactant ratio. Methanol substitution contributed something to the conversion

of nitric oxide into nitrogen dioxide and subsequently yielded more ozone than

the blank substitution case. However, the substitution of methanol and

isobutene (formaldehyde) produced a less reactive mixture and less ozone than

the baseline urban surrogate case.

The twenty-seven to one ppm carbon to NOx ratio results are presented in Fig.

VI.12. A very rapid conversion of nitric oxide into nitrogen dioxide was

observed for both the baseline and the methanol substitution cases. After

allowing for experimental uncertainties the ozone yield from the methanol

substitution experiment was found not to be significantly greater than that from

the baseline condition. Therefore, methanol substitution exhibited comparable

reactivity at this very high reactant ratio. The pronounced effect of ppm

carbon to NOx ratio on methanol's substitution reactivity is the subject of

further analysis and future experimental work.

VI.6 CONCLUSIONS

The chemical kinetic mechanism derived from this experimental evidence is likely

VI.1 9
DATE lAUG 26

DATE iAUG 11

DATE iAUG 18

TYPE t 8S BASE PPBC/NOx i 3.04

TYPE i 2/3 BASE PPBC/NOx i 1.98

TYPE s 8S + MS PPBC/NOx : 3.32

1200

NO

N02 1000

OZONE

NO

MS => MEOH SUBSTITUTION

BS »> BLANK SUBSTITUTION

8S -> BASELINE

600--

PARTS

PER

BILLION40o

200--

120 240 360

MINUTES OF IRRADIATION

480 600

FIGURE VI. 10 : BASELINE SURROGATES. 2/3 BASELINE & METHANOL SUBSTITUTION AT 3:1
1200-r

NO

NQ2 1000- -

OZONE

800 -f

600--

PARTS

PER

BILLION4QQ

200--

NO

N02

i \N f S

W

DATE iJUL 28

DATE rAUG 01

DATE : AUG 24

TYPE i 8S BASE

TYPE : 2/3 BASE

TYPE : 8S + MS

PPBC/NQx : 8.86

PPBC/NOx : 6.16

PPBC/NOx : 8.88

BS -> BLANK SUBSTITUTION

8S -> BASELINE

OZONE, 8S BASELINE

OZONE. MEOH SUBSTITUTION

OZONE, BLANK SUBSTITUTION

120 240 360

MINUTES OF IRRADIATION

480 600

FIGURE VI.11 : BASELINE SURROGATES. 2/3 BASELINE & METHANOL SUBSTITUTION AT 9:1
1200-r

PARTS

PER

BILLION4QQ

DATE i AUG 10

DATE «AUG 09

DATE :AUG 25

TYPE i 8S BASE

TYPE « 2/3 BASE

TYPE : 8S + MS

PPBC/NQx t 27.14

PPBC/NOx * 17.58

PPBC/NOx : 27.78

OZONE. BLANK SUBSTITUTION

OZONETTiEOH SUBSTITUTION

OZONE. 8S BASELINE

200

K., 120

— '"•-•••i

240 360

I I 1

480 600

MINUTES OF IRRADIATION

FIGURE VI. 12 : BASELINE SURROGATES. 2/3 BASELINE & METHANOL SUBSTITUTION AT 27s1
to show strongly beneficial effects when incorporated into urban EKMA-type

computer models. The reason is that, urban ppm carbon to NOx ratios in Los

Angeles, for example, are currently in the nine to one range (the evidence in

Fig. VI.11 is applicable). Existing plans and strategies for urban ozone

control emphasize increased control of hydrocarbon emissions. Hence, the

accelerated benefits of methanol sustitution at low ppm carbon to NOx ratios

appears very promising as an urban ozone control strategy.

VI.7 REFERENCES

VI.1. Sakamaki, F., S. Hatakeyama and Y. Akimoto; "Formation of Nitrous Acid

and Nitric Oxide in the heterogenous Dark Reaction of Nitrogen Dioxide

and Water Vapor in a Smog Chamber," Int. J. Chem. Kinet., V.15, P.1013-

1029; 1983.

VI.2. Whitten, G.Z., J.P. Killu s and H. Hogo, "Modelin g of Simulated

Photochemical Smog with Kinetic Mechanisms," Vol 1. Final Report, EPA600/3-80-028a, U.S. Enviornmental Protection Agency, Office of Research

and Deelopment, Research Traingle Park, North Carolina, 1980.

VI.3. Wu, C.H. and H. Niki, "Methods for Measuring N02 Photolysis Rate,"

Environ. Sci. Technol., Vol. 9; P.46; 1975.

VI.4. Schere, K.L. and K.L. Demerjian, "Calculation of Selected Photolytic Rate

Constants, over a Diurnal Range," EPA-600/4-77-015, U.S.E.P.A., 1977.

VI.23
1. Report No. 2. Government Accession No. 3. Recipient's Catalog No.

4. Title and Subtitle

Research and Development of Neat Alcohol Fuel Usage in Automobiles

5. Report Date

June 1984

6. Performing Organization Code

7. Autbor(s)

Pefley, R.K., Browning, L., Espanola, S., Pullman, B.,

Gururangan, R., Saito, N.

8. Performing Organization Report No.

ME84-1

10. Work Unit No.

9. Performing Organization Name and Address

The University of Santa Clara

Mechanical Engineering Research Department

Santa Clara, California 95053

11. Contract or Grant No.

NASA Lewis Contract NAG-3-143

13. Type of Report and Period Covered

12. Sponsoring Agency Name and Address

NASA Lewis Research Center

21000 Brookpark Road

Cleveland, OH 44135

14. Sponsoring Agency Code

15. Supplementary Notes

16. Abstract

The study is an extension of previous work. Its topical coverage ranges from research in

the forms of experimental operation of engines in cold chambers and a photochemical smog

chamber with allied computer modeling of the thermofluid,mechanics of cold engine starting

and photochemical smog through engine corrosion and wear studies, to vehicle operation,

which includes a description of the engineering development of after-market'conversion

kits to change Ford Pinto vehicles from gasoline to alcohol. Although further research

is recommended, evidence presented contributes positively to the view that it is feasible

to operate neat alcohol fleets in the range of 1000-10,000 vehicles, without major

operational difficulties and without exacerbating the environmental, health and safety

risks associated with petroleum fuels.

17. Key Words (Suggested by Author(s))

MetHanoi Emissions

Fuel Smog

Cold Start Corrosion and Wear

18. Distribution Statement

Approved for public release; distribution

unlimited

19. Security Classif. (of this report)

Unclassified

20. Security Classif. (of this page)

Unclassified

21. No. of Pages 22. Price"

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