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dissertation entitled

APPLICATION OF A CONCENTRIC VORTEX-CELL

BIOMASS FURNACE TO GRAIN DRYING
presented by

Eliud Ng'ang'a Mwaura

has been accepted towards fulfillment
of the requirements for

.Ph.D. degree in Agricultural Engineering

[7/3%//M

Major professor

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APPLICATION OF A CONCENTRIC VORTEX-CELL

 

BIOMASS FURNACE TO GRAIN DRYING

By

Eliud Ng'ang'a Mwaura

A DISSERTATION

Submitted to
Michigan State University
in partial fulfillment of the requirements
for the degree of

DOCTOR OF PHILOSOPHY

Department of Agricultural Engineering

1984

ABSTRACT
APPLICATION OF A CONCENTRIC VORTEX-CELL
BIOMASS FURNACE T0 GRAIN DRYING
By
Eliud Ng'ang'a Mwaura

Artificial crop drying is relatively energy intensive and uses
about 60% of the energy required to produce corn, exclusive of irriga-
tion. Biomass fuels, such as wood chips and corncobs, are potential
alternatives to fossil fuels for crop drying.

A concentric vortex-cell furnace (CVCF) with a maximum heat
output of 690 kw was built and incorporated in an in-bin counterflow

(IBCF) corn dryer with an hourly capacity of 3 m3

at an initial grain
temperature of 10°C, a drying temperature of 71°C, and corn dried from
25% to 15.5% moisture content (wet basis). The system was tested using
wood chips and corncobs. Experimental tests included: (1) fuel con-
sumption and drying capacity. and (2) corn contamination with polycyclic
aromatic hydrocarbons (PAHs) and heavy metals. The results were com-
pared with the performance of a propane (LP) fueled system.

The results demonstrate that the CVCF biomass furnace is a
technically viable alternative to conventional fossil fuel burners.
The drying capacity of the CVCF/IBCF system is approximately 90% of

the LP fueled system. The reduction in the drying capacity is a result

Eliud Ng'ang'a Mwaura

of the reduction in the flow rate of the drying air mass due to the
heating of the air before the fan entrance.

The CVCF/IBCF system operates at about 70% efficiency in con-
verting biomass fuels into energy for grain drying. The average
specific energy consumption (kJ/kg of water removed) of the system is
about 6,100 kJ/kg compared with 4.600 kJ/kg for the LP fueled system.

There is no objectionable discoloration or objectionable odor
in the corn dried with the CVCF/IBCF system. The grain contains no
dangerous concentration levels of PAH or heavy metals.

The operating costs (exclusive of labor and fixed costs) of the
CVCF/IBCF system are 30 to 40% of the energy costs of the LP fueled
system. The total break-even costs of the CVCF/IBCF system are about
20% higher than the total costs of the LP system (at 1983 prices) due
to the high capital investment required to convert the LP system into
a biomass system.

The CVCF furnace does not have immediate economic feasibility

in the USA at the 1983 fossil fuel prices and interest rates.

.:-I

. . _.=1:." "

 

To Kinnthi and his colleagues

'i'f'iitm), a“. .

f-. _

ACKNOWLEDGMENTS

The author expresses deep appreciation to Professor F. v.
Bakker-Arkema for his guidance and encouragement during the course
of this research. He has provided moral and technical support since
1979 as a teacher, mentor, and friend.

Sincere thanks are expressed to Dr. G. R. Van Ee (Agricultural
Engineering), Dr. J. F. Steffe (Agricultural Engineering), Dr. E. N.
Braselton (Animal Health Diagnostic Laboratory), Dr. E. Grulke (Chemi-
cal Engineering), and to Mr. S. J. Kalchik (Kalchik Farms, Bellaire,
MI) for serving as guidance committee members. Their advice, along
with that of Dr. A. Dhanak (Mechanical Engineering) and Dr. A. Atriya
(Mechanical Engineering), is greatly appreciated.

Special gratitude is sincerely expressed to the Netherlands
Government and the University of Nairobi, Kenya, for financial support
and scholarship, and to Michigan Department of Commerce (Energy
Administration), and Shivvers Corporation, Corydon, Iowa, for providing
funds for this project.

The author is indebted to Mr. and Mrs. Steven Kalchik and
Mr. and Mrs. Larry Kalchik, Bellaire, MI, on whose farm this research
was conducted. Their help and hospitality is sincerely acknowledged.

Two fonmer students who assisted in data collection during the
early days of this project deserve special mention: Juan Rodriguez

and Carlos Fontana. Their assistance is appreciated.

iii

Sincere thanks are extended to the entire community of
Agricultural Engineering Department and International Studies Center
for making the author's stay at Michigan State University a happy
one.

Special thanks goes to my family, Eva, Njoki, Hangari, and
Mumbi for their companionship, love, and encouragement.

The author can be reached at the following address:

Department of Agricultural Engineering
University of Nairobi (Kabete Campus)
Post Office Box 30197

NAIROBI2 KENYA

iv

 

TABLE OF CONTENTS

Page

LIST OF TABLES . viii

LIST OF FIGURES xi

LIST OF APPENDICES . . . . . . . . . . . . . . xiii

LIST OF SYMBOLS . xiv
Chapter

1. INTRODUCTION . . . . . . . . . . . . . . 1

1.1 Michigan Grain Production 3

1.2 Grain Preservation . . . . . . 4

1.3 Biomass Fuels for Grain Drying . . . . . 6

1.4 Grain Production and Storage in Kenya . 8

1.5 Objectives . . . . . . . . . 11

2. LITERATURE REVIEW . . . . . . . . . . . . 13

2.1

2.2

2.3

Biomass Furnaces for On- Farm Grain Drying. . . 13
2.1.1 Major Products of Combustion . . . . 14
2.1.2 Design Criteria for Biomass Furnaces . . 15
2.1.3 Direct- Combustion Biomass Furnaces . . . 16
2.1.4 Gasification-Combustion Furnaces . . . 24
2.1.5 Direct-combustion Versus Gasification-

Combustion Systems . . . . 27
2.1.6 Emissions from Biomass Furnaces and

Grain Contamination . . . . . . 30
Biomass Combustion Theory . . . 31
2.2.1 The Properties of Biomass Relevant to

Combustion . . . . . 32
2.2.2 The Process of Biomass Combustion . . . 45
Grain Drying . . . . . 65
2.3.1 Importance of Grain Drying . . . 66
2.3.2 Corn Quality as Affected by Drying

Methods . . . . . . . 67
2..33 Drying Systems . . . . . . . . . 74

Chapter Page
2.4 Grain Drying Theory and Simulation . . . . . 90
2.4

Thin- -Layer Drying Models . . . . 91
2.4.2 Equilibrium Moisture Content Models . . 95
2. 4. 3 Deep- bed Drying Models . . . . 96

3. FURNACE DESIGN, CONSTRUCTION, AND OPERATION . . . . 99

3.1 Design and Construction . . . . . . 99
3.2 Combustion Principle of the Concentric

Vortex- Cell Furnace . . . . . . . 103

3.3 Furnace Operation . . . . 104

3.4 Concentric Vortex- Cell Furnace Design Theory. . 104

3.4.1 The Design Heat Output. . . . 105

3.4.2 Grate Design . . . . . . . . . . 105

3.5 Fuel Feed System . . . . . . . . 107

3.5.1 Furnace Air Supply System . . . . . 112

4. THEORY . . . . . . . . . . . . . . . . 123

4.1 Furnace Perfonmance Simulation . . . . . . 123

4.1.1 Preliminary Estimates . . . . 124

4.1.2 Heat Transfer from Combustion Chamber . . 128
4.2 Heat Transfer Between the Furnace and the Hood . 136
4.3 Grain Drying Simulation . . . . . 140

5. EXPERIMENTAL INVESTIGATIONS . . . . . . . . . 143

5.1 Furnace/Dryer Performance . . . . . . . . 143
5.2 Chemical Analysis . . . 144
5.2.1 Determination of Corn Contamination with

H

PA .
5.2.2 Detennination of Corn Contamination with
Heavy Metals . . . . . . . . . . 148

6. RESULTS AND DISCUSSION . . . . . . . . . . . 149

6.1 Introduction . . . . . . . . . . . . 149
6. 2 Fuel Consumption . . . . 149
.2.1 Factors Affecting the Fuel Feed Rate . . 158

6. 2.2 Effect of Fuel Moisture on Fuel Consump-

ti on . . 162

6.2.3 Effect of Excess Air on Furnace Perform-
ance . . . . . . . . 166
6.3 In-Bin Counterflow Corn Drying . . . . . . 176
6.3.1 Biomass Fuel Feed Rate . . . 179
6.3.2 Experimental Specific Energy Consumption . 179
6.3.3 Drying Capacity . . . 181
6.3.4 Standardized Dryer Performance . . . . 181

vi

Chapter Page

6.4 Co orn Drying Simulation . . . . . . . . . 184
6.4.1 Cycle Times . . . . . . . . . . 184
6.4.2 Drying Capacity . . . . . . . 190
6.4.3 Specific Energy Consumption . . . . . 190
6.4.4 Airflow Rate . . 194
6.4.5 Comparison of Experimental and Simulated

Grain Moisture Contents . . 194
6.4.6 Comparison of Experimental and Simulated
‘ Drying Data . . . . . 196

6.5 Predicted Drying Performance Parameters . . . 198
6.5.1 Effect of Refill on Dryer Performance . . 200
6.5.2 Effect of Dryeration on System a

Performance . . . 202
.5.3 Effect of Drying Air Temperature on
System Performance . . . 203
6.5.4 Effect of Initial Moisture on System
Performance . . 204
6.5.5 Effect of Ambient Temperature on System
Perfonnance . . . 205

6.6 Economics of the CVCF/IBCF System . . . . . 205
6.6.1 Operating Costs . . . . . . . 206
6. 6. 2 Capital Budgeting Analysis . . . 209

6.7 Corn Contamination in Drying by Direct Biomass
Heating . . . . 213

6.8 General Observations . . . . . . . . . 217
6.8.1 Fuel Feed System . . . . . . . . . 217
6. 8. 2 Grate Performance . . . . . . . . 218
6. 8. 3 The Combustion Chamber . . . . . . . 219
6. 8. 4 Furnace- -to-Dryer Duct . . . . . . . 219
6.8.5 Safety . . . . . . . . . . 220
6. 8. 6 Miscellaneous . . . . . . . . . . 220

7. SUMMARY . . . . . . . . . . . . . . . . 221

8. CONCLUSIONS . . . . . . . . . . . . . . 224

9. SUGGESTIONS FOR FURTHER RESEARCH . . . . . . . 226
APPENDICES . . . . . . . . . . . . . . . . . 228
REFERENCES . . . . . . . . . . .

 

Table
1.1

1.2

2.1

2.2

2.3

2.4
2.5

5.1

6.1

6.2

6.3

6.4

LIST OF TABLES

World production (1981) of main cereal grains,
1,000 metric tons .

World production of soybeans and pulses (1981),
1,000 metric tons .

Proximate analysis of selected biomass fuels (weight,
percent, dry basis)

Ultimate analysis (percent dry basis) and higher heat-
ing values (MJ/kg) of selected fuels .

The effect of moisture content on heat recovery and
combustion efficiency of wood . .

Thermal conductivity of selected wood species

Numerical grades and sample grade requirements for
S corn . -

Excitation/emission wavelengths, retention times, and
detection limits of different PAHs on the fluores—
cence detector . . . .

The experimental and simulated fuel consumption rates
of the CVCF biomass furnace at different ambient
conditions and different drying air temperatures

Experimental and simulated airflow, drying energy,
total energy losses and furnace/system efficiencies
for tests 1- 7 given in Table 6.1 .

Effect of ambient air conditions on fuel consumption,
furnace and system efficiency, and total heat loss
from CVCF/IBCF drying system; dr rying air temp =

80° C, drying airflow = 12. 3 m3/m -m1n . . .

Furnace/IBCF system perfonmance using woodchips at

different moisture contents (percent wet basis) for
different drying air temperatures . . .

viii

Page

34

36

39
42

68

147

150

152

159

163

01

.10

O!

.11

The theoretical effect of excess combustion air
(percent) on the CVCF furnace perfonmance

Experimentally measured higher heating value (MJ/kg)
of wood chips and corncobs at different moisture
contents (percent wet basis) . . . . . .

Actual drying rates, energy (fuel) consumption rates
and operating costs (1983 prices) for seven drying
tests at different ambient conditions at the Kalchik
Farms, Ballaire, MI . . .

Standardized energy (fuel) consumption and operating
costs (1983 prices) for different fuels and fuel feed
rates for the CVCF- IBCF drying system . .

A comparison of simulated and experimental CVCF- IBCF
performance parameters for Test No. . . .

Simulated operating conditions during Test 1
Comparison of CVCF-IBCF experimental and simulated
results for seven tests under different drying
conditions . . . . . . . .
Simulated energy consumption, drying rate, and operat-
ing costs for different dryer operation modes at
average ambient conditions .
Estimates and assumptions for a 12-year budgeting
analysis (1983 prices) of the CVCF/IBCF system drying
in Mode 1 . .

Economic analysis results of different biomass fuels
compared to propane (1983 prices) in drying 380 tons
of corn annually from 26% to 15. 5% w. b. . .

Concentration of polycyclic aromatic hydrocarbons on
corn dried with biomass energy and LP- -gas

Experimental results: Test No. 1
Experimental results: Test No. 2
Experimental results: Test No. 3

4

Experimental results: Test No.

Page

168

177

178

182

185

186

197

201

210

212

215
232
233
234
235

 

 

Page
Experimental results: Test No. 5 . . . . . . . 236 i

     

Experimental results: Test No.6 . . . . . . . 237
7 ’_.;Experimental results: Test No. 7 . . . . . . . 238 '1':

Figure
2.1
2.2
2.3
2.4
2.5
2.6
2.7
2.8
2.9
2.10
2.11

3.1
3.2

3.3
4.1

4.2
4.3
4.4

LIST OF FIGURES

The Cell Furnace .

Vortex Furnace

European Sloping Grate Principle

Schematic Diagram of Husk Fired Cyclone Furnace
Up-draft Gasification and Close—coupled Combustion
Down-draft Gasification and Close-coupled Combustion .
Batch Column Dryer

Schematic of a Low-temperature In—bin Drying System
Schematic of a Crossflow Grain Dryer . . . . . .
Schematic of a Concurrent Flow Dryer

Schematic of the Internal View of an In-bin Counter—
flow Drying "Shivvers System“ . .

Schematic of the Concentric Vortex-cell Furnace

Schematic Diagram of Magnetic and Bin Switches for
Fuel Level Control in the Fuel Hopper . .

The Schematic Diagram of the Eductor

Schematic Flow Diagram of the MSU Concentric Vortex-
cell Furnace . . . . .

Simplified Furnace Diagram
Radiation-Convection Heat Transfer . . . . . . .

Complete Radiation Network for the System 1,2, 3, and
4 in Figure 4. 2 . . . . .

xi

Page
17
19
21
23
26
28
77
79
83
85

87
100

110
117

125
131
133

135

 

Figure

4.5

6.1
6.2

6.3

6.4
6.5

6.6

Radiation Network for the Surfaces 3, 5, and 8 in
Figure 4. 2 . .

Fuel Feed Rate as a Function of Fuel Moisture

The Effect of Excess Air on the Combustion Chamber
and in the Exit Flue Gas Temperature .

The Effect of Excess Air on the Drying Air Tempera-

ture, Furnace Efficiency, and the System Efficiencey .

Cycle Times vs Number of Cycles for Test No. 1 .

Simulated Specific Energy Consumption as a Function
of Number of Cycles . . . . .

Simulated Airflow Rates at the Beginning of Each
Cycle

xii

Page

137
164

169

170
188

193

195

 

LIST OF APPENDICES

Appendix
A. Sample of CVCF Computer Input/Output Data
B Experimental Results
C. CVCF Computer Program
0

In-Bin Counterflow Drying Computer Program

xiii

Page
229
231
239
258

 

LIST OF SYMBOLS

Area, m2

Ash content, percent

Inert heat absorbing compound
Concentration kg/m3

Convective thermal resistance 9C/N
Specific heat kJ/kg °C

Diameter m

Diffusivity constant mZ/hr
Diameter m

Activation energy kJ

Radiated thermal energy kJ/h

View factor for radiation from surface m to surface n
Friction factor

Fractional void volume

Mass of wet grain kg

Grashoff number

Gravitational constant 9.8 kgm m/kgfs2
Enthalpy kJ/kg

Absolute humidity kg/kg

Higher heating value kJ/kg

Heat of combustion kJ

Radiocity w/m2

xiv

 

Ra
RH
rh
59
St
SECO

N><£<C¢+—-|

Thermal conductivity

Low heating value kJ/kg

Length m

Moisture content percent wet basis

Mass flow rate kg/h

Moisture content decimal dry basis

Mass fraction

Pressure kPa

Prandtl number

Rate of heat transfer kJ/h

Gas constant

Radius m

Rayleigh number

Relative humidity percent

Relative humidity decimal

Specific gravity

Stanton number

Specific energy consumption kJ/kg of water
Temperature °C

Time h

Velocity m/s

Volume m3
Height of water removed kg
Mole fraction

Excess air percent

XV

 

p

0'

Emissivity

Efficiency percent

Grain temperature °C

Density kg/m3

Stefan-Boltzmann constant 5.669 x IO'BH/mzk

Subscripts

a
B

C

Air

Bin

Combustion products

Convective

Dried grain

Eductor nozzle exit

Equilibrium

Furnace combustion chamber

Evaporation

Increment

Increment

Moisture

Increment

Induced mass flow at initial conditions
Products of combustion

Dried product, i.e., grain

Reactants in a combustion process
Radiation heat loss

Motive fluid at initial supply conditions

Sky
xvi

 

 

.oP
T

 

Vapor

Hall

water

 

 

CHAPTER 1

INTRODUCTION

Cereal and pulse grains are a major source of food for human
and animal consumption. In terms of world production in 1982, wheat
ranks first, followed by corn, rice, and barley (Table 1.1). As a
source of human food, rice has no equal among the grains, especially
in the Orient where the major (91%) of the world's supply of rice is
produced (FAO, 1982). Corn, millet, and sorghum are the main cereal
grains for human consumption in Africa. According to USDA (1982), the
USA produced in 1981 about 46% and 20% of the total world production
of corn and all the cereals, respectively. Other important cereal
grains are barley, oats, and rye.

Pulse grains, such as beans and peas, are a significant source
of protein for human consumption, especially in Third World countries
where meat consumption is comparatively low. Pulse grains are pro-
duced mainly in Asia (52% of world production), Africa, and South
America (see Table 1.2). Soybeans are an important source of oil and
protein for both human and animal consumption. The USA is the major
producer of soybeans and produces more than 60% of the total world

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Table 1.2.--World production of soybeans and pulses (1981), 1,000
metric tons.

 

 

Region Soybeans Pulses
Africa 319 5,076
Asia 10,320 22,241
Europe 552 2,401
S. America 19,575 3,294
U.S.A. 55,260 1,673
World 87,941 42,403

 

 

 

 

 

Sources: FAO (1982), USDA (1982).

1.1 Michigan Grain Production

Michigan is a major grain producing state. Among the major
grain producing states in the USA, Michigan, in 1982, ranked first
in production of dry beans (30.0%), 6th in oats (4.6%), 8th in corn
(3.7%), 7th in rye (3.1%), 22nd in wheat (1.2%), and 18th in soybeans
(1.4%) (MDA, 1983). The shelled corn annual production in Michigan
increased from 4.94 in 1978 to 7.81 million tons in 1981. This repre-
sents an increase of 58%. During the same period, annual soybean
production increased from 0.55 to 0.82 million tons, an increase of
49%. Annual wheat production showed the highest increase from 0.42
million tons in 1978 to 1.13 million tons in 1981, an increase of
153%. Other grains showed production increases of 20% (dry beans)

to 31% (barley); the oats production decreased by 10% (MDA. 1983).

1.2 Grain Preservation

Grains are stored either on the farm or in commercial storage
facilities. Storage structures should provide suitable conditions for
maintaining the grain quality. The stored grain should be protected
from insects, rodents, molds, and the weather elements. Before stor-
age, it should be properly conditioned.

Grain drying is the most widely used method of preserving
grain (Brooker et al., 1974). Moisture is removed to prevent the
development of a favorable environment for the growth of molds and
insects. Grain drying may be either natural or artificial. Natural
drying occurs in the field before harvest, in an ear corn crib, or
when on the ground during solar drying.

Artificial drying is an energy intensive system. According to
Brooker et al. (1974), about 60% of the energy required to produce corn
exclusive of irrigation is used for artificial drying. In Spite of
the high energy demand for grain drying, more than 80% of the corn
produced in the USA is artificially dried (ASAE, 1978). Artificial
drying has the following advantages:

a. allows early harvest which, in turn, reduces storm and
shattering field-losses, and permits early land prepa-
ration for the next crop;

b. allows planning of the harvest season to make better
use of the labor;

c. allows long-time storage without deterioration;

d. enables fanners to take advantage of higher prices

at some time after harvest;

 

e. maintains grain viability and quality;
f. allows the use of full season hybrids (longer
maturing varieties) that yield more grain per
hectare.
Considering the above advantages, it is obvious that grain drying
cannot be omitted without undue loss of grain quantity and quality at
the farm level.

The current artificial grain dryers use convection-type
heaters that burn fossil fuels, usually liquefied petroleum (LP) gas
or natural gas. According to Friedrich (1978), high-temperature
dryers operate at about 50% efficiency inutilizing fuel to evaporate
moisture from grains. It takes from 4650 to 8140 kJ to remove one
kilogram of water from grains such as corn, depending on the initial
and final moisture content, the amount of fines, the dryer design,
and the weather conditions (Brooker et al., 1978).

It is obvious that the current high-temperature dryers require
large amounts of energy. Also, the fossil fuels compete with domestic
and industrial uses. It is, therefore, desirable to develop drying
systems which can reduce fossil fuel requirements for grain drying.

There are four principle possibilities of reducing fossil
fuel usage (ASAE, 1978):

1. increased use of high moisture corn storage to

feed animals at farms close to where corn is grown;

2. use of natural air or low temperature drying;

 

3. use of combination drying systems employing the
advantages of high-speed, high—temperature drying and
low-speed, low-temperature (or natural air) drying;

4. use of solar heated drying systems;

5. application of crop residues as a crop drying
energy source.

The last item is the subject of the research in this dissertation.

1.3 Biomass Fuels for Grain Drying

Biomass energy is a form of solar energy stored by plants in
the form of hydrocarbons. Many researchers have investigated utiliza-
tion of this energy (Steffgen, 1974; Shrader, 1977; Hall, 1979; Stout,
1979; Claar et al., 1980a and 1980b; Claar, 1981; Anderson et al.,
1981; Sukup et al., 1982; Morey et al., 1980 and 1982; Richey et al.,
1980). Steffgen (1974) observed that the 540 million tons of agri-
cultural wastes produced in 1970 in the USA, if converted to fuel oil,
could have replaced more than half the fossil fuels needed to generate
electricity, or could have replaced all the natural gas and fuel oil
used that year for generating electricity. Burwell (1978) noted
that 303 million tons of cereal straw and corn stalks/cobs could have

9 GJ of energy used by the USA in

supplied about 5.5% of the 71.x10
1975.

In making these estimates, Steffgen (1974) and Burwell (1978)
did not account for the residues that are needed to conserve soil
fertility. If this factor is accounted for, the total energy avail-

able is much less than the above estimates.

 

Agricultural biomass can be converted to oil by liquefaction,
to gas by hydrogasification or pyrolysis, and to alcohol by fermanta—
tion. It can also be directly converted to thermal energy by combus-
tion. The combustible energy value of biomass is relatively high and
has, therefore, been a subject of considerable research (Pimentel et
al., 1981; Helsel and Wedin, 1983).

Before discovery of fossil fuels, wood was the principal fuel
for mankind. As recently as 1900, wood accounted for 25% of the total
US energy supply; by 1976 it provided less than 1.5% (Fisher, 1976).
By comparison, 8% of Sweden's and 15% of Finland's energy needs were
met by wood in 1977 (Merril and Gage, 1978). Biomass is still the
major source of energy in Third World countries, especially in rural
areas.

One ton of shelled corn produces about 0.25 tons of corncobs.
Depending on the efficiency of combustion and the design of the dry-
ing system, corncobs from 0.25 ton of shelled corn can dry one ton of
corn from 26 to 15.5% wet basis* (Richey et al., 1980). Thus, if all
the corncobs are collected, they can be converted to enough energy to
dry the entire US corn crop and also provide thermal energy for other
farm applications such as heating of buildings and processing of other
crops.

During periods of low grain prices, the grain may have a greater

value as a fuel than as a feed or food. 0f the corn crop, 3 to 5%

 

*All moisture in this thesis are stated on a wet basis unless
otherwise indicated. '

can provide enough energy to dry the remainder of the crop with an
efficiently designed combustion system (Morey and Thimsen, 1980).

The forest products industry derives 45% of its total energy
need from burning bark and mill wastes. Much of the wood waste is not
utilized and is thus available to nearby farms. It may be also prac-
tical to set aside some land for biomass fuel production. Commercial
forests for fuel is another possibility (Stobaugh and Yergin, 1980).

Incineration of biomass for crop drying requires a well designed
furnace to ensure complete combustion. Incomplete combustion may cause
undesirable odors; discoloration and contamination of grain (Anderson
et al., 1981; HUtt and Winkler, 1978; Jacko et al., 1982; Joe at al.,
1982). Thus, development of efficient and clean burning solid fuel
combustion systems is an important step in increasing the utilization

of crop and residue materials for heat energy.

1.4 Grain Production and Storage in Kenya

Cereal grains, in particular corn and legumes, constitute the
staple food for most Africans, including Kenyans. Cereals are grown
by "small scale" (under 20 hectares) and "large scale" (over 20 ha)
farmers under different ecological, social-economic, environmental and
technological conditions. The production problems are many and varied.

Large scale farmers produce grains almost entirely as cash
crops and usually have the equipment and the know-how to handle the
crops more efficiently than the small scale farmers. The latter,
depending on how “small" they are, produce grains for home consumption

and for cash if they have a surplus. Some grain is consumed fresh

 

before it matures and the rest is stored for seed and for future con-
sumption. Small scale farmers face major problems in drying, cleaning,
and storage of grains.

Handling practices vary from place to place within Kenya. Some
farmers, especially the small scale ones, harvest unhusked ear corn and
suspend it from trees or place it over fires in their homes or on
special platforms. The latter method is an example of utilization of
biomass energy for crop drying, especially along the coastal region of
Kenya. In other regions traditional granaries are found in which
unhusked or dehusked ear corn is kept for natural drying. The granaries
vary in size, in accessability for inspection, in cleaning and removal
facilities, and in the degree of protection they offer from insects,
rodents, and the weather. Some corn is sun-dried (either as ear-corn
or shelled grain) by spreading it on mats, or on the ground.

Solar drying may be economical, but is risky due to uncertain-
ties in the weather, and in the labor availability for handling. There-
fore, most farmers leave the crops in the field to dry naturally, and
harvest when the grain is sufficiently dry (16 to 18%) for solar
drying. This results in field infestation by insects and molds, and
consumption by birds and rodents. Also, the land preparation for the
next crop is delayed. To avoid the delay, some farmers cut the corn
stalks and stack them in 2 to 4 meter diameter conical stacks and leave
them in the field to dry for up to four months.

Long-tenn storage of grain is undertaken by the Kenya Cerals

Board (a Kenya government agency). The Board is in charge of drying

10

and storage of cereal grains for the purpose of price stabilization
and for emergency needs (such as occur during a drought). In 1981 the
Board handled about 20% of the total Kenyan corn production of 500,000
metric tons. The remainder is handled by individual farmers or coop-
eratives. The Board buys grain from farmers at a specified standard
quality and government price.

For long-term strategic storage (up to 3 years), the Board
uses Cyprus bins. A Cyprus bin is an oval-shaped concrete structure
which is constructed halfway underground. Corn is dried to 12% before
it is placed in a Cyprus bin. The capacity of a typical Cyprus bin is
about 250 tons. For short-term storage, corn is stored at 13% mois-
ture in 90 kg bags in masonry sheds.

According to FAO (1982), the Kenyan corn production was 1.45
million metric tons in 1979; in 1980 it was 1.75 million tons and in
1981 2.20 million tons. The average annual grain losses on small—
scale farms were estimated at 16%. This amounts to a monetary loss in
Kenya of approximately $42 million annually (MAO, 1981). There are
also health hazards caused by aflatoxin contained in moldy grain
(Anderson and Pfost, 1978).

Most of the grain losses in Kenya occur during storage. Insect
damage is the most severe problem. Artificial drying is essential in
the surplus-producing areas with high rainfall. Long-term storage
conditions are most favorable at the higher altitudes due to the low
temperatures and low relative humidities. Judging from the available

literature on grain drying and storage in Kenya, it appears that the

 

11

current storage facilities and practices do not provide the necessary
grain protection from insects and molds. Since the most severe insect
and mold infestation occurs in the field, it is reasonable to assume
that earlier harvesting and artificial drying can contribute substan-
tially in minimizing the losses.

Kenya has no fossil fuel or mineral resources. It is, there-
fore, necessary to design grain drying systems that are effective,
inexpensive, energy (fossil fuel) efficient and made of locally
available materials. Such systems must utilize renewable biomass
energy sources, such as crop residues and wood. Since biomass waste
and wood are also used for domestic energy needs, there is a special
need in Kenya to design efficient biomass combustion furnaces for

grain drying.

1.5 Objectives

The general objective of this research is to design a direct
fired vortex-cell biomass furnace and incorporate the unit in an
existing on-farm grain drying system. The specific objectives are:

1. To develop, build, and test a prototype of a direct-
fired combustion, concentric vortex-cell biomass
furnace for on-farm grain drying.

2. To demonstrate the feasibility of the furnace as a
technical alternative to LP gas and natural gas
burners for on-farm grain drying.

3. To integrate the furnace into a commercial in-bin

counterflow drying system.

 

12

To test commonly available (in Michigan) renewable
biomass fuel sources such as wood chips and corncobs.
To determine the safety aspects of the biomass furnace
with respect to chemical abSorption of smoke con-
stituents by the grain.

To detennine the economics of drying with biomass energy.
To model semicontinuous in-bin counterflow grain
drying.

To model a concentric vortex-cell biomass furnace.

 

 

CHAPTER 2
LITERATURE REVIEW

2.1 Biomass Furnaces for On-Farm Grain Drying

Since the early 70's many investigations have been conducted
on the use of biomass energy for grain drying and other farmstead
operations. Most research has been directed towards the study of
direct combustion systems for converting crop residues and processing
wastes into thermal energy. Two main categories of combustion systems
are found in the literature: direct-combusion systems, and
gasification-combustion systems.

Studies on collection and direct-combustion of cobs and corn
stover have been conducted at Iowa State University since 1975
(Kajewski et al., 1977; Claar et al., 1980b; Anderson et al., 1981;
Wahby et al., 1981; Sukup et al., 1982; Anderson et al., 1983). A
direct-combustion incinerator-type furnace burning cobs was developed
by Pioneer Hi-Bred for drying ear seed-corn (Dahlberg, 1977). Seed
Grains, Inc., also developed a corncob-fired furance (Claar et al.,
1980b).

Crop residue burners have been marketed by several companies
in the USA (Moraczewski, 1980), in Germany (Strehler, 1976), and
Canada (Peill, 1980).

In all the furnaces referred to above, biomass is converted

into thermal energy by direct-combustion.{{1n direct-combustion
K

13

 

14

furnaces there is no distinct separation of the pyrolysis and combus-
tion processes. Both processes occur in the same physical combustion
chamber.

Gasification of biomass to produce producer gas, which is
subsequently burned to provide energy for crop drying, has been studied
at the University of Minnesota (Morey et al., 1980; Morey et al., 1982);
at Purdue University (Foster et al., 1982); at Clemson University
(Payne et al., 1981); at the University of Florida (Shaw et al., 1982);
and at the University of Kentucky (Payne et al., 1980). No commercial

units are, as yet, being marketed in the USA.

2.1.1 Major Products of Combustion

The major products of combustion of biomass fuels are heat,
exhaust gases, and ash; the products of incomplete combustion of the
fuels are tar and char.

The amount of heat released in combustion depends upon the
heating value of the fuel. The heating value is determined by measur-
ing the enthalpy change between reactants (fuel and oxygen in air)
and products at 25°C in an adiabatic bomb calorimeter. The heating
value is reported as the higher (gross) heating value (HHV), or the
lower (net) heating value (LHV). The HHV represents the heat of
combustion if the water vapor produced is condensed to a liquid at
25°C. The LHV is the heat of combustion when the water remains as
vapor. The LHV is lower than the HHV because the latent heat of
vaporization is not recovered. Heating values are reported in both
wet and dry fuel basis. The unit of the heating values is kJ/kg of

fuel.

 

 

15

The gaseous products of biomass combustion include C02, H20,
CO, N2, and traces of N0 and S02. If the oxygen level is low, the
gases may contain tar. (lag is a mixture of volatile hydrocarbon
liquids (or vapors at elevated temperatures) produced by destructive
distillation of biomass.) The gases appear as smoke if soot is present.
S99; consists of a fine dispersion of black particles, chiefly carbon,
produced by theincomplete combustion of hydrocarbon fuels. Thus,
smoke is a vaporous mixture containing soot (and other unburned parti-
cles), water and tar vapors, and the other gases resulting from incom-
plete combustion.

Char (charcoal) is a black carbonaceous solid residue left
after most of the volatile compounds, such as tar and water vapor, have
been distilled from biomass fuels as a result of incomplete combustion
or anaerobic heating. The char can be burned in oxygen releasing
heat, C02, and ash.

Ash is the uncombustible residue (after complete combustion)
consisting of a mixture of inorganic compounds. At high temperatures
(above 750°C) the ash melts into a liquid, which, upon cooling solidi-
fies into a glassy solid called slag (or clinker). Slag is undesir-

able in furnaces because it may clog grates and air supply openings.

2.1.2 Design Criteria for Biomass Furnaces

Claar et al. (1981) considered the following factors to be of
significance in the design of a biomass furnace:

1. The furnace should be able to supply heat for several

farmstead applications.

 

 

16

2. The furnace should be capable of handling several
biomass fuel types.

The furnace system should be of simple construction.
The furance should require minimum labor

The furnace should be easy to control and maintain.

0'1th

The furnace should provide favorable combustion
characteristics.

The last requirement (6) is the most difficult to achieve and
the most critical. Incomplete combustion may cause contamination (with
hazardous compounds) and discoloration of the grain when flue gases
are used directly for drying the grain. If a heat exchanger is

used, it can become clogged by products of incomplete combustion.

2.1.3 Direct-Combustion Biomass Furnaces

There are four main types of direct-combustion furnaces
employed in grain drying systems: (1) the cell furnace, (2) the vortex
furnace, (3) the sloping grate furnace, and (4) the cyclone furnace.
The design features and performance characteristics of these furnaces

are discussed in the following sections.

2.1.3.1 The Cell Furnace

The cell furnace was originally developed to burn sugarcane
bagasse, and later modified to burn wood. The furnace consists of a
firebrick-lined cylinder with primary and secondary air supplied
through tangential ports in the wall (see Figure 2.1). Buchele et al.

(1981) have described the principle of operation of the cell furnace.

 

 

EXHAUST

 

 

CYCLONIC
ACTION

 

 

 

 

Figure 2.1. The Cell Furnace.

 

 

 

18

The distinguishing feature of the cell furnace is that fuel
enters the furnace at the point of the greatest turbulence, usually
near the top of the furnace. Fine fuel particles are separated from
the heavier ones and are burned in suspension. The heavier particles
fall onto the fuel pipe on the hearth at the bottom of the cell. Pri—
mary air is introduced to the pile of fuel in the hearth through
many small ports in the refractory wall. The air and heat rapidly
penetrate the pile (since there are no fine particles), and rapid
drying and combustion takes place. The middle of the furnace consists
of a restricted throat around which secondary combustion air is
introduced through tangential ports producing a swirling turbulent
action that intimately mixes the fuel and the air.

The cell furnace issuitable for burning chopped fuels such as
wood chips and corncobs. The heat release is about 5 GJ/h-m2 of grate

area (Claar et al., 1981).

2.1.3.2 The Vortex Furnace

The vortex-type furnace has been a subject of considerable
research at Iowa State University (Buchele et al., 1981; Wahby et al.,
1981; Claar et al., 1981; Anderson et al., 1983). The vortex—type
furnace consists of two chambers, one above the other. The lower
chamber dehydrates and gasifies thepfile of biomass which may be
placed on a solid hearth or on a grate (see Figure 2.2). The char
remaining after gasification is burned in the lower chamber. The

volatile gases are completely burned in the upper chamber.

Steam General ion

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Figure 2.2. Vortex Furnace (from Claar et al., 1981).

20

To enhance complete combustion, air is introduced in the com-
bustion chamber through tuyeres located tangentially in the wall of
the furance. This creates a vortex spiral which causes the flame to
have a turbulent and tight spiral. Thus, any unburned material or fly
ash is centrifugally separated from the flame spiral, and thrown back
onto the burning pile.

The wall of the furnace is lined with firebricks for insula-
tion. The fuel is fed onto the hearth (or grate) by a side-mounted
feed auger. The heat output and the flue gas temperature are con-
trolled by regulating the fuel feed rate and the combustion air flow
rate.

The Iowa State University (Claar et al., 1981) furnace is
fabricated from two concentric steel cylinders. The inner cylinder
is lined with firebricks. Combustion air is blown into the space
between the two cylinders, and thus, preheated before entering the
combustion chambers.

The heat release capacity of vortex furnaces is 5—9 GJ/h-m2 of
grate area depending on the overall combustion efficiency, the fuel
type, the feed rate, the condition of the fuel, and the insulation

of the furnace (Claar et al., 1981; Kenzler et al., 1982).

2.1.3.3 The Sloping Grate Furnace

The sloping grate furnace consists of a firebrick-lined com-
bustion chamber (see Figure 2-3). The fire hearth is made of an
inclined grate which supports the fuel during drying and combustion.

The fuel is continuously introduced into the furnace at the top part

 

 

21

 

 

BURN our
ARCH
35°
31°
DRYING FUEL GASIFICATION ----
AND BURNING FINAL BURN-OUT ——0

AND ASKING
FIXED, BARE TUBE RECIPROCATING DUMP
SLOPING GRATE FOOT CRATE CRATE

 

 

 

 

Figure 2.3. European Sloping Grate Principle (MacCallum, 1979).

 

 

22

of the grate. The fuel is dehydrated as it passes over the upper
section. It then descends into the lower burning section. The ash
is continuously removed from the lowest part of the grate. About
75-80% of combustion air is intrdouced under the grate (MacCallum,
1979); the 20—25% is introduced through tuyeres located at the burn-
out arch.

The sloping grate furnace is only suitable for boilers with
heat exchangers (due to a fly ash problem). It has been modified for
crop drying by introducing extra secondary combustion air above the
grate (as in a vortex furnace) and by minimizing the undergrate primary
combustion air (Buchele et al., 1981). The furnace can burn relatively
high density wet (SS-70% wb) fuels such as wood chips and corncobs.
Straw and husks are not suitable since they are blown off the grate
before complete combustion

Heat release of large sloping grate furnaces used in boilers
is about 8-10 GJ/h-m2 depending on the fuel used and its moisture

content (MacCallum, 1979).

2.1.3.4 The Cyclone Furnace

A cyclone-type furnace has been developed in India (Singh et
al., 1980). The furnace was designed to burn rice husks and generates
heat for heating grain drying air, or for generating steam (in a boiler
for parboiling paddy in the rice mill).

The cyclone furnace consists of a combustion chamber and a
cyclone chamber (see Figure 2.4). The combustion chamber is fabricated

from an oil barrel (55x140 cm) with a 40 degree conical bottom. An

 

 

.IMIl-Hzgii’

Figure 2.4.

23

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(Singh et al., 1980).

 

24

exit pipe at the cone bottom provides the outlet for the flue gases.
The ashes are trapped at the bottom of the cone and are continuously
removed.

The cyclone chamber consists of a cylinder (40 cm diameter and
70 cm long) mounted on the side of the combustion chamber and inclined
at an angle of 30°. The rice husks are blown tangentially into the
cyclone chamber by an air blower. The husks are thus kept rotating
throughout the combustion chamber. The husk-air mixture is ignited by
an oil burner.

The cyclone furnace is suitable for burning rice husks, chopped
straw and other finely ground materials. The heat release is similar

to that of the vortex furance (5-9 GJ/h-mz).

2.1.4 Gasification-Combustion Furnaces

The direct heating of dryers with the exhaust flue gases from
biomass combustion requires clean and complete burning to avoid
contamination of the grain. Direct-combustion of biomass materials
may result in smoke and fly ash. Gasification of the biomass to pro-
ducer gas is a viable alternative to direct burning.

Richey et al. (1980) listed four advantages of gasification
compared to direct-combustion furnaces:

1. Minimum pollution of air without the use of heat

exchangers.
2. More efficient conversion of biomass to heat energy

(BO-90% of the energy is recovered).

 

 

25

3. Low combustion temperatures (650-850°C) compared to
direct-combustion (1000-1100°C), thus reducing the
cost of construction materials and also reducing the
formation of slag.

4. Easier control of the combustion rate (by regulating

primary airflow).

Biomass is gasified by burning the material with limited air
to produce "producer" gas consisting of C0, H2, and some CH4. The gas
contains impurities such as C02, water vapor, and nitrogen. The heat-
ing value is 37,452 kJ/m3 compared to 37,260 kJ/m3 for natural gas
(Richey et al., 1980).

The major experience with biomass gasification has been gained
during World War II when West European countries could not obtain
petroleum products. Portable gasifiers using charcoal and wood chips
were used to power motor vehicles and tractors. These units had power
output equal to approximately 75% of their gasoline rating (SERI, 1979).

It is possible toburn producer gas in a conventional dryer
burner without visible smoke or particulates. However, it is difficult
to avoid condensation of tar as the gas cools (Payne et al., 1980).
Therefore, most investigators report that the gas should be burned in
a combustion chamber attached to the gasifier unit (Richey et al.,
1980; Payne et al., 1980, 1981; McCoy et al., 1981; Morey et al.,
1980, 1981; Richey et al., 1981).

Gasifiers are either of the updraft or downdraft type. In the

updraft design (Figure 2.5), solid fuel is intrdouced at the top of

26

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EXHAUST

     
   

    

 
 

  

  
     
 

 

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Figure 2.5. Up-draft Gasification and Close-coupled Combustion.

 

 

27

the gasifier and is supported on a grate at the bottom. The fuel is
ignited at the bottom of the pile on the grate. Air is forced upward
through the grate, supporting partial combustion on the grate. The
producer gas is collected at the top. The unburned fuel acts as a
filter for particulates entrained within the gas. Updraft gasifiers
of up to 9.5 MJ/h have been developed by seed companies to utilize
corncobs as a fuel for drying the corn (link at al., 1982).

In downdraft gasifier systems, the air is blown from the top
(or the sides) of the fuel chamber as shown in Figure 2.6. Combustion
takes place in the funnel shaped bottom with the pyrolysis gases passing
down and out through thehigh-temperature char. The high temperature
cracks the tars into stable gases. The system is more suitable for
internal combustion engines (Richey et al., 1981) and for grain drying
systems in which the producer gas is burned in a gas burner (similar
to a propane gas burner).

In cross-draft gasifier systems, the air intake is through a
horizontal nozzle in one side near the bottom of the chamber. The gas
exists through a vertical grate in the opposite side. The burning zone
is adjacent to the air nozzle (Richey, 1980).

Srivastava and Posselius (1981) applied dimensional and simili-
tude modeling to develop scaling laws for biomass gasifiers. Their

techniques can be applied to direct combustion furnaces.

2.1.5 Direct-combustion Versus Gasification-Combustion Systems
Judging from the literature available, no comparative study

has been conducted on the engineering performance and/or the economic

EXHAUST

28

 

 

   

i" .:

    

  

 

9.0.:0'07" O‘N'E'I: .-. .‘el- 9
Wopi‘fflm?

   
 

 

   
 
 

   

SECONDARY AIR D—D

Down-draft Gasification and Close-coupled Combustion.

Figure 2.6.

29

feasibility of direct-combustion and gasification-combustion systems
for grain drying. It has been postulated that gasification systems
have several advantages over direct-combustion. The advantages include
(1) minimum air pollution due to low particulate emissions and complete
combustion, (2) more efficient conversion of biomass to heat, (3) lower
combustion temperature and thus, lower construction cost, and (4) more
effective control of combustion rate. According to a study by Barrett
et al. (1981), there is no significant difference between heat recovery
efficiency.and particulate emission from a direct-combustion (concen-
trix vortex-cell) furnace and two types of gasification-combustion
furnaces (a down- and an up-draft furnaces).

Gasification is not absolutely essential for crop drying
installations when the heat is required in situ. The difference
between the direct-combustion and the close-coupled gasification-
combustion furnaces is the physical distance between the gasification
and combustion chambers. In the direct-combustion furnace, biomass is
dehydrated and gasified at the bottom of the furnace; secondary com-
bustion of the gases starts immediately above the gasification section
and continues throughout the secondary chamber. In the close-coupled
gasification-combustion process the gases are burned in a second com-
bustion chamber attached to the gasification chamber.

A well designed concentric vortex-cell furnace is more effi-
cient than a gasification system since there is less heat loss between
the gasification and combustion sections. For internal combustion
engines and for replacing natural or LP gas systems for domestic or

other similar uses, gasification systems are superior.

30

2.1.6 Emissions from Biomass Furnaces and Grain Contamination

During the drying process the products of combustion are
usually introduced directly into the grain. Some compounds may be
retained by the grain through condensation or absorption.

The relevant combustion products are C02, CO, H20, aldehydes,
ketones, polycyclic (or polynuclear) aromatic hydrocarbons (PAH), $02,
503, N0, N2, soot and suspensions of heavy metals. The soot and other
particulates absorb high quantities of the other combustion compounds
including PAH (Winkler, 1977, cited by Hfitt et al., 1978). The quan-
tity and specific types of combustion compounds depend on the fuel
used, and on the completeness of the combustion process. Solid fuels,
which have not undergone a coking process, have long carbon chains and
thus, produce more hydrocarbons than gaseous fuels (Hfitt et al., 1978).

The PAH are the most critical compounds. Souci (1968) and
Hangebrauck et al. (1964) reported that benzo [a] pyrene, benz [a]
anthracene, dibenzo [a,h] anthracene, benzo [a] anthracene and benzo
[a] pyrene have shown some carcinogenic activity in test animals.

Hfitt et al. (1978) identified 15 PAH compounds on corn dried with
direct heaters using different commercial fuels. Contamination levels
were on the order of 0 to 3 parts per billion (ppb) for well operating
fuel burners. Maladjusted burners resulted in considerably higher
concentrations. More recently, Joe et al. (1982) detected 11 PAH com-
pounds in artificially dried barley malt. The PAH concentrations in
the malt varied from 0.1 to 1.5 ppb (except for phenanthrene which
showed a concentration of 2.1 ppb). Hfitt and Winkler (1978) detected

a benzo [a] pyrene (BaP) concentration of 0 to 0.61 ppb in 21 samples

31

of freshly harvested corn; they concluded that corn may contain trace
PAHs prior to drying when grown near major highways or industrial
plants.

Most of the grain dried in the USA is used for livestock feed.
Thus, the risks to human beings is minimal. According to the Deutsche
Forschungsgemeinschaft (1976), the extent to which PAHs from contami-
nated animal feeds are transferred into meat for human consumption is
not clear.

According to Lijinsky and Ross (1967), a charcoal broiled
steak may contain 50 ppb BaP, depending on the temperature and the type
of fuel used in the broiler. Smoked foods contain from trace up to 6.6
ppb of BaP (Santodonato et al., 1980). Thus, it can be concluded that
smoked and charcoal broiled foods may contain higher levels of carcino-
genic and other PAH compounds than grain dried with well-designed and
well-operated direct-heated biomass fueled grain dryers.

Jacko and Barrett (1981) described the equipment and procedures
for determining particulate concentrations and size distributions in
the exhaust gas streams of biomass furnaces. Barrett et al. (1981)
compared the particulate emissions in biomass furnaces with those in
LP gas burners; average emission was 0.19 kg/h from biomass furnaces,

and 0.00028 kg/h from LP gas burners.

2.2 Biomass Combustion Theory
Biomass is any material derived from growing organisms. The
energy contained in the biomass is stored in plants by the photosynthe-

sis process. Combustion of biomass materials is one of the

32

bioconversion processes whereby the biomass energy is transformed into
usable energy. All biomass materials found on farms such as wood,
grain, straw, corncobs, etc., are potential sources of energy. Wood
and corncobs are the most commonly used biomass fuels due to their
favorable combustion and transportation characteristics (Buchele et
al., 1981).

The theory on biomass combustion is related to the theory of
combustion of coal and wood. It should be noted that the coal
combustion mechanism is different from that of wood due to the differ-
ent aging processes and high volatile component contents of wood and

coal.

2.2.1 The Properties of Biomass Relevant to Combustion

The structure, physical and thermal properties, and chemical
composition of biomass materials influence their utility as a fuel
feedstock. The literature on the properties of wood and corncobs will
be reviewed since these materials are the principle fuels used as

energy sources for on-farm crop drying.

2.2.1.1 Chemical Properties

Combustion is an oxidation process by which carbon is oxidized
to carbon dioxide, and hydrogen is oxidized to water. Combustion of
biomass materials consists of a series of chain reactions. The exact_
pathways of the reactions are determined by the structure of the
combustible molecules. The structure determines the location and

accessability of the carbon and the hydrogen by the oxygen. Also, the

33

higher fuel heating value is affected by the carbon content of the
fuel (Graboski and Bain, 1979).

The major chemical constituents of wood and corncobs are
cellulose, hemicellulose and lignin (Tillman et al., 1981).

The proportions of the constituents depend upon the species of the
wood. Generally, softwoods have 40-45, 24-37, and 25-30% (dry basis)
cellulose, hemicellulose and lignin, respectively. Hardwoods contain
40-50, 22-40, and 10-38 percent of these compounds, respectively
(Shafizadeh and DeGroot, 1976). Thus, hardwoods generally have a
higher hemicellulose and a lower lignin content and sometimes a
slightly higher cellulose content; the cellulose percentage in both
types of wood is similar.

Foley and Hooven (1981) determined the proportions of cellu-
lose, hemicellulose, and lignin in corncobs. Their analysis included
other compounds such as xylan, pectins, and hexogan, and also the dis-
tribution of these compounds in the woody and pith/chaff portions of
the corncobs. Corncobs contain (on the average) 40, 36, and 6 percent
cellulose, hemicellulose and lignin, respectively, on a dry basis.

The exact proportions of these compounds differ among the different
varieties of corn.

In addition to the cellulose, hemicellulose and lignin, wood
and corncobs contain extractable compounds such as volatile oils
(terpenes and sesquiterpenes), resins and fatty acids, pigments, and
water soluble carbohydrates (starch, simple sugars, and organic acids).
Wood and corncobs with a high extractive content have higher heating

values (see Table 2-1) (Tillman et al., 1981). The nature and quantity

34

Table 2.1.--Proximate analysis of selected biomass fuels (weight,
percent, dry basis).

 

 

 

 

 

 

 

 

Fuel Volatile matter Fixed carbon Ash Source
Wood
Cedar 77.0 21.0 2.0 a, b
Douglas Fir 86.2 13.7 0.1 a, b
Hemlock 84.8 15.0 0.2 a, b
Bark
Cedar 86.7 13.1 0.2 a, b
Douglas Fir 73.0 25.8 1.2 a, b
Hemlock 74.3 24.0 1.7 a, b
Husklage 59.4 37.8 2.9 c
Corncobs 80.1 18.3 1.6 d
80.3 18.1 1.6 e
Coal 6.4 81.4 12.2 f
a = Howlett and Gamache (1977).
b = Mingle and Boubel (1968).

c a Sukup (1982).

lb 0.
ll ll

-15
ll

Ebeling and Jenkins (1983).
Payne et al. (1980)
Babcock and Wilcox (1978).

 

35

of extractables determine the products after pyrolysis and gasification

(Graboski and Bain, 1979).

2.2.1.1.1 Proximate and Ultimate Analysis

The proximate and ultimate analysis are terms expressing the
chemical-physical characteristics and the chemical composition of a
fuel without reference to the physical form in which the compounds

appear. The proximate analysis classifies the fuel in terms of its

 

moisture, volatile matter, ash, and fixed carbon contents. Thus, the
proximate analysis provides information on the volatility of the fuel.
Compared to coal, wood has a larger percentage of volatile matter and

a smaller percentage of ash (Tillman, l980). The proximate analysis of
wood and corncobs has been published by several authors (Babcock and
Wilcox, 1978; Hewlett and Gamache, 1977; Thimsen and Morey, 1981;

Claar et al., 1978; Ebeling and Jenkins, 1983). Table 2.1 presents the
proximate analysis for selected biomass fuels.

The ultimate analysis contains the elemental analysis of a

 

fuel. Generally, the percentages of carbon, hydrogen, nitrogen,
sulfur, oxygen, ash, and heating values are reported. The composition
of wood and corncob fuels is fairly constant. These fuels do not
contain sulfur and have very little nitrogen (Claar et al., 1978).
Table 2.2 presents the ultimate analysis for selected biomass fuels.
The free moisture (above the chemically bound water) in solid fuels
results in erroneously high values of hydrogen and oxygen in the

ultimate analysis because moisture is detected as additional hydrogen

36

Table 2.2.--Ultimate analysis (percent dry basis) and higher heating
values (Md/kg) of selected fuels.

 

 

 

Fuel C H 0 N S Ash HHV Source
Wood
Cedar 48.8 6.4 44.5 -- -- 0.37 18.0 a
Douglas Fir 52.3 6.3 40.5 0.1 -- 0.08 21.1 a
Hemlock 50.4 5.8 41.4 0.1 0.1 2.20 20.1 a
Average 50.5 6.2 42.1 0.1 -- 0.88 19.7 -
Bark
Douglas Fir 56.2 5.9 36.7 -- -- 1.20 22.1 a
Corncobs 46.7 5.8 44.8 0.5 0.2 1.60 18.7 b
47.6 6.0 44.2 0.5 0.1 1.63 18.6 c
46.4 6.1 44.0 0.3 0.1 2.85 18.6 d
46.6 5.9 47.1 0.4 -- 1.60 -- f
Average 46.8 6.0 45.0 0.4 0.1 1.92 18.7 -
Coal 82.1 2.3 2.0 0.8 0.6 12.20 30.8 e

 

 

 

 

 

 

 

 

 

ID . D. O U" 9’
ii II II II II

‘9'!
ll

Tillman (1978).

Payne et al. (1980).
Sukup (1982).

Ebeling and Jenkins (1983).

Babcock and Wilcox (1978).
Thimsen and Morey (1981).

 

37

and oxygen. Thus, the ultimate analysis is performed on dehydrated
fuel and reported on a dry basis.

The carbon content of biomass materials is lower than of coal
while the atomic carbon to hydrogen ratio is much higher in coals than
in biomass materials (Graboski and Bain, 1979). The bound oxygen
content of biomass materials is considerably higher due to alcohol
groups in the cellulose, hemicellulose, and lignin fractions of the
biomass (Kirk and 0thmer, 1963). Thus, biomass materials have lower
heating values than coal. The heating values of most biomass materials
is in the range of 15-25 MJ/kg of biomass dry matter.

As stated previously, the nitrogen and sulfur contents of bio-
mass are considerably lower than of coal. Therefore, sulfur and
nitrogen pollutants resulting from direct-combustion of biomass are
small enough to meet EPA standards (Graboski and Bain, 1979).

The ultimate analysis can be used to calculate the empirical
chemical formula of a fuel. The formula is useful in developing com-
bustion equations, calculating the stoichiometric air required for
combustion, and predicting the airborne emissions. The fuel higher
heating value can also be predicted using the ultimate analysis.
Several authors have proposed empirical equations for predicting the
higher heating value of biomass fuels using the ultimate analysis:

Tillman (1980):

HHV = 0.475C - 2.38 (2.1)

38
Dulong-Bertholot (Spiers, 1962):
HHV = 0.475c + 1.44H - 2.9 * 10’4(N + 0) + 0.0935 (2.2)
Institute of Gas Technology (1978):
HHV = 0.34C + 1.32H + 0.069 - 0.12(N + O) - 0.015As (2.3)

Where HHV is the higher heating value in MJ/kg dry matter, and C, H, N,
O, and As are the percentages (dry basis) of carbon, hydrogen, nitro-
gen, oxygen, and ash as given by the ultimate analysis of the fuel.

According to Graboski and Bain (1979), equation (2.3) is
preferred for estimating higher heating values.

The proximate and ultimate analyses are determined using the
American Society for Testing Materials (ASTM) Standard numbers
0-317'—73,D-3175-77, and 0-3174-73 for the proximate analysis; and
numbers 0-3178-73, 0-3177-75, and 0-2361-66 for the ultimate analysis;
number 0-3286-77 for the higher heating value; and number 0-3683-78 for
the ash analysis (Institute of Gas Technology, 1978).

2.2.1.2 Biomass Physical and Thermal Properties
The combustion mechanisms (the chemical reaction process as
determined by the reactants, their kinetics, chain reactions, and

equilibrium) of biomass materials depend upon their physical and

thermal properties and in addition upon their composition and heating
values. The relevant properties are: moisture content. thermal
conductivity, specific heat, heat capacity, density (as related to

void volume), and particle size distribution.

39

The moisture content of a biomass fuel is a most important
property because it affects the net heating value of the fuel, the
heat recovery, the combustion efficiency, and the quality of the flue
gases. The other physical and thermal properties, and to some extent
the chemical properties, are also influenced by the moisture content
(MC). Table 2.3 shows the effect of wood MC on the heat recovery and

combustion efficiency.

Table 2.3.--The effect of moisture content on heat recovery and
combustion efficiency of wood.a

 

 

 

Moisture content Recoverable heatb Combustion efficiency
(‘30 (NJ/kgl (5’4)
0.0 16.51 82.5
4.8 16.37 81.8

13.0 16.08 80.4
20.0 15.80 78.9
28.6 15.36 76.8
42.9 14.37 ' 71.8
50.0 13.65 i 68.2
60.0 12.22 61.1
71.4 9.35 46.7

 

 

 

 

aFrom Bliss and Black (1977).

bTheoretical values based on a maximum heating value of 20 MJ/kg,
an initial wood and air temperature of 17.0°C, and 50% excess air.

Wood, corncobs, and other biomass materials are hygroscopic.
They absorb and desorb water depending upon the moisture content, the

relative humidity and the temperature of the surrounding air, and the

40

inherent moisture retention capacity. Skaar (1972) and the USDA
Forest Products Laboratory (1974) reported the maximum moisture con-
tents of different wood species in the heartwood and sapwood parts of
a tree. Softwoods contain 23% to 46% and 52 to 71% maximum moisture
in the heartwood and sapwood, respectively. The maximum moisture
contents in the hardwoods are 31 to 62% in the heartwood, and 30 to 59%
in the sapwood. The data of Skaar (1972) indicate that the differ—
ence in moisture content between heartwood and sapwood is quite pro-
nounced in the case of softwoods, but is less significant in hard-
woods. The moisture content is higher in the sapwood than in the
heartwood in the case of softwoods. For hardwood the MC of the heart-
wood can be higher than that of the sapwood in some trees. The mois-
ture content of sapwood varies depending on the region and site of the
tree, while that of the heartwood is fairly constant throughout the
year due to the restricted movement of water. After a tree is har-
vested and chipped, its moisture content is in equilibirum with the
environmental conditions. .

According to the data of Foley and Hooven (1981), the MC of
the woody portion of corncobs is higher than the pith/chaff portion.
The moisture content of corncobs varies among varieties and the
harvest corn kernal moisture content.

Tillman et al. (1981) proposed the following equations for

predicting the energy required for drying wood:

<11 = (1-M)(100-Tf)(cd)(wf) (2.4)

41

qz =M(100-Tf)(cw)(w.) (2.5)
Q. =M(H.g)(wf) (2.6)
q“ =M<Hw)(w.) (2.7)
cd = [1.11 + 4.82*1o'3(1ave-273)] (2.8)
Qtota1 = <11 + £12 + 93 + q“ (2 9)

Where M is the moisture content of the wood (decimal); Tf the tempera-
ture of the fuel (K); Cd the specific heat of dry wood (kJ/kg°C);
Cw the specific heat of water (4.174 kJ/kg°C); Wf the weight of wet
wood (kg); Hfg the heat of evaporization ow water (kJ/kg°C); Hw the
heat of wetting (kJ/kg H20); Tave the average between the initial and
the final wood temperature (K) and Qtotal the total energy needed to
evaporate water from wood (kJ/kg).

The total energy for evaporating water from maple wood at 50%
MC and initially at 25°C is 1.36 MJ/kg if the wood is heated to a final
temperature of 100°C.

Stamm and Loughborough (1935) determined the heat of wetting
of wood (35.8-14.3 kJ/kg at a MC of 4-12 percent and 9.5-2.4 kJ/kg
at a MC of 12-30 percent).

Knowledge of the thermal conductivity of wood, corncobs, and
other biomass materials is necessary for modeling pyrolysis, gasifi-
cation, and corrbustion. The thermal conductivity is a function of

temperature, spatial direction, and the major constituents including

42

moisture, cellulose, hemicellulose, and lignin. The available thermal

conductivity data for wood are given in Table 2.4.

Table 2.4.--Thermal conductivity of selected wood species.

 

 

 

Wood type Temperature °C Conductivity W/m°C
Cypress, across grain -- 0.097
Pine, with grain 21.1 0.346
Oak, across grain 15.0 0.208
Fir, with grain 23.9 0.109
Maple, across grain 50.0 0.190
Pine board 43.3 0.102

 

 

 

 

Sources: cypress: Handbook of Chemistry and Physics (1966).
pine, oak, and maple: McAdams (1954); Gray et al. (1960).
fir: Chapman (1974).

As can be seen in Table 2.4, the available thermal conductivity
data are limited to the 15 to 50°C range. Thermal conductivity data at
higher temperatures at which pyrolysis and gasification occur are
lacking in the literature. Moreover, the available data do not
include the moisture ranges for which the thermal conductivity values
were measured. Thus, if thermal conductivity values are to be used
in modeling pyrolysis or gasification during combustion, reliable
data over the actual range of combustion conditions must be devel-

oped.

43

Several specific heat equations for wood at temperatures up
to 100°C have been proposed. Beall (1968), Wenzel (1970), and Skaar
(1972) developed the following equations:

Beall (1968): (2.10)

3

cf = 1.085 + 0.409*M + 2.535*10' *1 + 5.45*10'4*M*T (2.10)

Wenzel (1970):

3

Cd = 1.11 + 4.82*10- *T (2.11)
Skaar (1972):
Cf = Cd)1.-M) + 4.19*M (2.12)

Where Cf is the specific heat of wet wood (kJ/kg°C); Cd is the spe-
cific heat of dry fuel (kJ/kg°C); M is the moisture content (decimal);
and T is the wood temperature (°C).

As with the thermal conductivity, no references were found for
the specific heat and heat capacity of biomass materials for the
temperature range needed for pyrolysis and combustion. Hearmon and
Burcham (1955), cited by Mohsenin (1980), suggested that the specific
heat of wood, with or without moisture, does not depend greatly on
species.

Equations (2.10) and (2.12) do not give the same values of the
specific heat of wood, given the same temperature and moisture. For

wood at 50°C and 20% moisture, equation (2.10) gives the specific

44

value of 1.3 kJ/kg°C; equation (2.12) gives a specific value of
1.9 kJ/kg°C at the same condition.

The density of biomass fuels is required for determining the
energy content of the fuels on a volume basis. There are three ways
of defining the density of biomass. Bulk density refers to the
weight of a mass of material packed in a given container divided by
the container volume. Apparent (or unit) density is the weight of
each unit of the material divided by the volume of the unit; the pore
space within each of the unit is included. True (or solid) density
refers to the weight per unit volume of the solids within each unit
of the material. All three are used to report the density of biomass
materials. The density of biomass depends upon the nature of the mate-
rial, its moisture content and degree of compaction. Unprocessed
oven-dry wood, with 7 to 8 percent moisture, has an apparent density
of 641 kg/m3 (hardwoods) and 449 kg/m3 (softwoods). High moisture
content wood can have an apparent density as high as 961 kg/m3
(Graboski and Bain, 1979).

According to Tillman et al. (1981), the density of wood varies
among different trees of the same species and also within the same
tree. It increases from the pith to the bark, decreases with increas-
ing height, and varies within the same growth ring.

The apparent density of wood, corncobs, and other biomass
materials varies with moisture content. Graboski and Bain (1979)
proposed the following relationship between the moisture content M

(decimal wet basis, as given by the biomass proximate analysis), and

45

the dry and the wet biomass apparent densities, pad and par’

respectively:

pad = (1.-ll)par (2.13)

The thermaland the physical properties of most farm wastes
including corncobs. corn husks, and wood chips are unavailable in the
literature. Due to the inherent nature of biomass, their properties

are characterized by a wide variability.

2.2.2 The Process of Biomass Combustion

Biomass materials contain a variety of chemical constituents,
some of which are not combustible (inert), such as water and inorganic
ash. The combustible constituents include volatile hydrocarbons and
nonvolatile carbonaceous and polymeric compounds. The combustion
mechanism and the heat and emissions generated in a biomass furnace,
depend upon the nature and the proportions of the chemical constituents
of the fuel.

The range of combustible biomass materials is very diverse with
respect to fuel composition and the state of aggregation. Not only is
the chemical composition of biomass fuels variable, but large hetero-
geneities are usually encountered within the same type of fuel. The
particle size and the thermophysical properties also vary between
different materials and within samples from the same material. The
wide range of chemical composition and physical pr0perties complicates
the theoretical analysis of the combustion process of biomass fuels

(Edwards, 1974).

46

The combustion process of wood has been studied more exten-
sively than of any other biomass fuel (Junge, 1975; Tillman, 1980;
Shafizadeh and Chin, 1977; Tuttle, 1978; Browne, 1958). It will be
reviewed to illustrate the combustion processes for biomass materials.

Wood combustion involves the oxidation of the carbon and
hydrogen to carbon dioxide (002) and water (H20), as shown by the

following equations:

C + 02 + C02 + 32.8 MJ/kg of C (2.14)
Hz + 1/202 + H20 + 141.9 MJ/kg of H2 (2.15)

Equations (2.14) and (2.15) are simplifications of a complex series of
pathways (involving free radical reactions) in which solids are con-
verted to gases and other solids through solid-phase pyrolysis. Also,
the gaseous compounds are converted into radicals by decomposition.
The radicals react with each other and with the oxygen in the com-
bustion air (Browne, 1958). While the CO2 and H20 are the major

final products. there are numerous intermediate radicals and com-
pounds such as OH, CH2, CZHS’ H00, CO, etc.

Edwards (1974), and Shafizardeh and DeGroot (1977) simplified
wood combustion assuming that the process is characterized by a series
of distinct events (or reaction phases): (1) solid phase precombus-
tion reactions, (2) gas-phase reactions, and (3) char oxidation.

The above events and the chemical pathways occurring within
them, determine the rate and degree of completion of reactions (2.14)

and 2.15), and also the formation of pollutants such as the oxides of

47

nitrogen and polycylic aromatic hydrocarbons (USDA Forest Service,

1982; Chakraborty and Long, 1968).

2.2.2.1 Solid-phase Precombustion Reactions

The precombustion reactions for wood (and for most biomass
solid fuels) are the initial processes occurring on the grate of a
biomass furnace. The precombustion reactions are heating, evaporation

of moisture. and pyrolysis of the dehydrated solid.

2.2.2.1.1 Heating and Drying

Before initiation of the pyrolysis reactions, the water in the
solid fuel must be evaporated and the solids heated to the pyrolysis
temperature (SOD-625°C) (Edwards, 1974; Browne, 1958). The heating
and drying are endothermic processes. The specific heat of wet fuel
is a function of the moisture content and may be estimated by equation
(2.12) (Skaar, 1972). Thus. some of the heat of combustion is used
to heat and dry the fuel.

The water in the fuel may limit the maximum temperature of the
core of individual fuel particles. The heat transfer rate from the
surface to the center of the particle depends upon the thermal conduc-
tivity of the fuel particles. The thermal conductivity of wood and
other biomass products increases with moisture content as indicated
by the following equation proposed by the US Forest Products Labora-
tory (1974):

48

7?
II

[(5.18 + .096*H)Sg + .57*Fvv] x 4.18 *10‘4 (2.16)
where: thermal conductivity, W/cm-K
mositure content, decimal

$9 = specific gravity

Fvv = the fractional void volume, decimal

37?
ll

Thus, moisture increases the rate of heat transfer from the surface to
the center of a fuel particle. However, no overall drying or heating
rate can be given for wood (and other biomass fuels) since particle
sizes and individual particle moisture contents vary dramatically
during high-temperature heating (USDA Forest Service, 1982). Also,
drying occurs in a very short time, possibly in seconds due to the high

combustion temperatures (Tillman et al., 1981).

2.2.2.1.2 Solid-phase Pyrolysis

Pyrolysis is an essential reaction in the solid fuel combustion
of basic biomass. Pyrolysis of carbonaceous materials is defined as
an incomplete thermal degradation, generally in the absence of oxygen
resulting in char, condensable liquids or tars and gaseous products
(Milne. 1979). Thus, pyrolysis involves the conversion of basic
biomass structures such as cellulose and lignin into more readily
oxidizable compounds. Slow pyrolysis occurs when materials are heated
in the absence of oxygen at temperatures above 200°C. In biomass
furnaces, rapid pyrolysis takes place in the presence of limited oxygen
(air) at about 1100°C. Browne (1958) discussed the theoretical course

of events in wood combustion preceding pyrolsis. Pyrolysis with and

49

without oxygen (air) can be presented by the following expression

(USDA Forest Service, 1982):

DRY W000 9-002 + CH4 + 02H6 + C3H6 + CH3COOH

+ CH3CH0 + TARS + CHAR (2.17)

Equation (2.17) is a simplification of the complex endothermic and
exothermic chemical chain reactions whose pathways and rates depend
upon the wood components, the particle size and geometry, the pyroly-
sis temperature and the heating rate. and the available amount of
oxygen. Browne (1958) reported pyrolysis temperatures of 200 to 500°C.
and 150 to 400°C in anaerobic and aerobic wood heating, respectively.

The pyrolysis of wood (and other biomass materials) is closely
related to the three major components of biomass, namely, the hemi-
cellulose. cellulose, and lignin (Milne, 1979; Antal et al., 1979;
Shafizadeh and Chin, 1977). The pyrolysis of biomass has been studied
by experimental investigations using each of the three components
separately. Shafizadeh and Chin (1977) reported the following
pyrolysis temperatures: hemicellulose 220-500°C, cellulose 320-380°C,
and lignin 220—500°C. Thus, in pyrolysis the hemicellulose, cellulose,
and the lignin fractions react differently.

Holocellulose is the total carbohydrate of wood, which is repre-

 

sented by cellulose and hemicellulose fractions. According to several
studies on pyrolysis of pure holocellulose and lignin (Shafizadeh and
Chin, 1977; Connors et al., 1980; Shafizadeh and Lai, 1972; Liu et al.,
1977; Brink, 1976; Goheen et al., 1976; Prahacs et al., 1971), char,

50

tar, water, 002, CH4, and 02H6 are the main products of pyrolysis. The
proportions of these products are different for each major component.
Lignin yields more char than holocellulose; the latter yields more tar
than the former. Examples of the exact products of pyrolysis can be
found in the previously cited literature (Fletcher and Harris, 1952;
Brink et al., 1973; Fang and McGinnis, 1976; Shafizadeh, 1968).

Milne (1979) suggested that it is reasonable to assume that
the pyrolysis of wood is closely related to the three major components
of biomass. Several other investigators reached this conclusion
although there has not been a quantitative demonstration of the rela-
tionship (Antal et al., 1979).

The products from wood pyrolysis are more complex than those
from the three major components of wood when pyrolyzed separately.
According to Milne (1979), no study has demonstrated that the products
of wood are the sum of the products of its components under comparable
conditions. Moreover, most pyrolysis studies are conducted in nearly
complete exclusion of oxygen. In biomass furnaces, a limited supply
of oxygen is furnished to sustain primary combustion. The behavior
and the products of pyrolysis under these conditions are not well
documented.

The kinetics of wood pyrolysis has been studied by several
researchers such as Kanury (1972), Kung (1972), Roberts (1970), and
Antal et al. (1979). Most kinetic analyses of pyrolysis are based
on heating samples at low temperatures (ZOO-400°C) without oxygen.
Thermal decomposition curves are fit using a general equation of the

form (Roberts, 1970; Milne, 1979):

51

dV n

a” = W (2.18)
where: k = Ae('E/RT) (2.19)

33-9 = (o-oflK (2.20)
where: K = kpe('E/RT) (2.21)

V and p are the fractions of total volatiles and the local
density of pyrolysing material, respectively; pf is the final density
of the pyrolysing material. In the above equations, the activation
energy E, and the product factors A, k, and n are not known with
adequate precision to predict aerobic pyrolysis at high temperatures in

biomass furnaces.

2.2.2.1.3 Influence of Moisture on Wood Pyrolysis

The moisture content of biomass fuel reduces the pyrolysis rate
and influences the amount of char and volatiles produced (Shafizadeh
and Lai, 1972). Wiggins and Krieger (1980) showed that there exists a
sharp physical separation between the solid-phase heating/drying zone
and the pyrolysis zone. This separation results from the temperature
gradient within the pyrolysing solid. Moisture reduces the flame
temperature in the furnace. thus, reducing the amount of volatiles
and increasing the amount of char produced. This results in a reduc-
tion of the rate of pyrolysis (Tillman et al., 1981).

The influences of moisture on the flame temperature can be
quantified by the relationship between moisture content, excess air and

the adiabatic flame temperature (Tillman et al., 1981):

52

T = 1920 - 151*Md - 5.15*Z (2.22)

ad

where Tad is the adiabatic flame temperature (K); Md is the moisture

(decimal dry basis); and Z is the excess air (percent).

2.2.2.2 Gas-Phase Oxidation
The volatiles formed in solid pyrolysis are oxidized via a
series of free radical pathways. Edwards (1974) discussed three zones
of gas-phase reactions:
1. The precombustion zone where chain initiation is
dominant;
2. The primary combustion zone where chain branching is
most prominent;
3. The post-combustion zone where chain termination

reactions occur.

2.2.2.2.1 Precombustion Reactions
Edwards (1974) showed the details of the pathways of the

precombustion reactions:

Csz + B +-26H3 + B (2.23)
CH3 + 02H6 + CH4 + (22H5 (2.24)
C2H5 + B + H + 02H4 + B (2.25)

H + C2H6 + H2 + CZH5 (2.26)

53

The compound 8 is any heat-removing molecule or particle such as ash
which is catalyst to the chemical reactions.

Precombustion gas-phase reactions are also considered pyroly-
sis reactions. The acetic acid and acetaldehyde may be degraded by
decarboxylation and decarbonylation as illustrated in equations

(2.27) and (2.28) (USDA Forest Service, 1982):

CH3C00H + CH4 + C02 (2.27)

CH3CH0 + CH4 + CO (2.28)

The above anaerobic reactions are predominant if the precom-
bustion zone is spatially increased by use of substoichiometric
quantities of undergrate air and high placement of overfire tuyeres as

is the case in a concentric vortex-cell furance.

2.2.2.3 Primary Combustion Reactions

When the gaseous products of pyrolysis are mixed with oxygen
in the primary combustion zone, they undergo a series of radical
reactions ultimately producing C02 and H20. The main reaction com-

pounds form chain branches (Hay, 1974):

CH3 + 02 + B + CH302 + B (2.29)
CH302 + CHZO + 0H - (2.30)

The above sequence of reactions generates the hydroxyl radical

as well as the main combustion intermediate CHZO. The latter reacts

54

with OH to form HC0 and CO. CH20 is the key intermediate resulting
from oxidation via the OH and oxygen attack (Tillman et al., 1981).
The concentration of CHZO reaches its maximum in flames at 1320 K
(which is within the range of wood and corncob combustion (Palmer,

1974).

2.2.2.3.1 Postcombustion Reactions
Chain termination reactions can be depicted by the following

representative sequence of reactions (Palmer, 1974):

HCO + 0H + C0 + H20 (2.31)
CD + OH +C02 + H (2.32)
CD + 02 -> CO2 + 0 (2.33)
0H + H + B + H20 + B (2.34)
CO+0+B+COZ+B (2.35)

The equilibrium and rate constants for these individual reactions have
been published (Jensen and Jones, 1978). However, no overall expres-
sions are available due to the many variables and complexities

associated with the oxidation of wood volatiles (Tillman, 1980).

2.2.2.4 Char Combustion
Bradbury and Shafizadeh (1980) described the empirical formula
for cellulose char as C6 7H3 30. The combustion of char is similar to

the combustion of carbon described by Kanury (1975):

55

1. Diffusion of oxygen to the fuel surface;
2. Absorption of the oxygen on the surface;
3. Reaction of oxygen within the solid to form
absorbed products;
4. Desorption of the absorbed products from the
surface;
5. Diffusion of desorbed products away from the
surface.
The slowest (H’ the above steps determines the burning rate. In the
case of carbon and char, combustion steps (2) and (4) are extremely
fast. According to Glassman (1977), char oxidation is considered to
be mass transport limited. This is due to the large particles and
relatively high temperatures associated with char combustion. Thus,
step (3) is much faster than steps (1) and (5). The burning rate is,
therefore, controlled by the diffusion rate of oxygen to the char
particles.
Several simultaneous char combustion mechanisms have been
suggested. Mulcahy and Young (1957) proposed the following mechan-

isms based on hydroxy radicals:

20H + C + C0 + H20 (2.36)

0H + C + 00 + H (2.37)
The classic Boudouard reaction was proposed by Glassman (1977):

C + C02 + 2C0 (2.38)

56
Bradbury and Shafizadeh (1980) proposed:

C* + 02 + C(O)m + C0 + CO (2.39)

2

C* + 02 + C(0)S + CD + 002 (2.40)

The asterick designates an active reaction site, the subscript m is a
mobile species and the subscript s is a stable species.

All the solid carbon, including the remains on the grate and
in the fly ash, must be oxidized via the above reactions to form C02.
The reactions must have sufficient residence time, high temperature, and
turbulance (the tree Ts) to complete the reactions in order to elim-

inate particulate emissions from the furnace.

2.2.2.5 Concentric Vortex-Cell Furnace Combustion Theory

The design and the operating principles of the concentric
vortex-cell biomass furnace are discussed in section 2.1.3.2 The
CVCF is similar to the spreader-stoker furnace if primary combustion
air is introduced into the fuel bed through the grate supporting the
fuel. The difference between a spreader-stoker and CVCF is that in the
former furnace, most of the combustion air is introduced into the
furnace through the grate. In the CVCF, secondary combustion air is
supplied tangentially through a set of tuyeres located above the fuel
bed. In some CVCF designs. the hearth is solid and no undergrate air
is supplied (Claar et al., 1980; Wahby et al., 1981).

Adams (1979, 1980) has developed a theoretical model for a

spreader-stoker wood-waste-fired boiler that predicts the flight times

57

and mass reduction of combustible particles entrained in the furnace
exhaust gases. The model also predicts the influence of fuel and
operating parameters on particulate emissions. The spreader-stoker
furnace employs the underfeed and overfeed principles described by
Nicholls and Eilers (1934). Thus, the analyzed furnace is similar to
the CVCF with undergrate air. Claar et al. (1981) suggested that this
model can be modified to analytically model the CVCF combustion
mechanism. However, the Adams (1979. 1980) models have not been
validated.

Several authors (Claar et al., 1980; Tuttle, 1977; Sukup,
1982) stressed the need to maintain high combustion temperatures and
turbulence, and to provide sufficient time for complete combustion.
The CVCF design provides suitable conditions for fuel combustion
(Claar et al., 1980). However, there has not been a study corre-
lating the combustion rate and the amount of particulate emissions from
the CVCF to the temperature. turbulence, and time aspects of the
CVCF. Thus, concentric vortex-cell furnaces have been designed and
built by trial and error, rather than by optimization of the three Ts
for a given furnace size and capacity (Claar et al., 1980 and 1981;

and Wahby et al., 1981).

2.2.2.6 Heat Transfer in Furnaces

Heat losses from a furnace determine the efficiency of fuel
conversion into useful energy. The net energy loss from a furnace at
steady-state conditions may be determined by the application of the

first law of thermodynamics. The conservation of energy equation for

58

a steady state flow, open system with one inlet and one outlet is

(Holman, 1980):

where:

Q

Ah

Qout

in

ri

pi
AHR

out =

2 2
Win + H1 - H2 + (U2 - U1)/29c (2.41)
n n
= 1.Elllmith. - AHR - iEIMPlAhPl (2.42)
hT - ho

net heat loss (kJ/h)
net work done on the system (kJ/h)

enthalpy of products entering (1),
or leaving (2) the system

enthalpy at temperature T(K) (kJ/kg)
enthalpy at 298 K (standard reference temperature)

mass flow rate (kg/h)

reactant i (i = air, fuel dry matter,

H20 in fuel, etc.)

product i of combustion (i = C02, H20, etc.)
heat of combustion at a reference temperature
of 298 K (kJ/kg)

gravitational constant (9.81 kg-m/kg $2).

velocity (m/s)

59

The heat loss determination using equations (2.41) and (2.42)
requires that the temperature and the composition of the flue gases are
known. 'Hueheat losses in the different components of the furnace can-
not be determined from equation (2.41).

Heat transfer in furnaces occurs in a combination of the three
modes of heat transfer, namely: radiation, convection, and conduction.
Heat transfer by convection, radiation, and conduction can be analyzed
using standard techniques available in numerous heat transfer books
such as Holman (1981) and Siegel and Howell (1981). However, heat
transfer by radiation and convection in flames within the combustion
chamber requires special treatment due to the thermal properties of
flames. Also, the combined radiation-convection heat transfer should

be studied since it occurs in a furnace.

2.2.2.6.] Convective Heat Transfer from Flames

Calculation of heat transfer in flames involves the determina-
tion of the convective heat transfer coefficient for the products of
combustion. The thermal properties of flue gases and particulates
entrained therein are quite different from those of air. Moreover,
the use of the arithmetic mean of the surface and gas temperatures
as the film temperature is inadequate due to the large variation in
the physical and thermal properties with temperature. Also. chemical
recombination reactions occur when the hot gases are cooled; consequently
there may be some heat release within the boundary layer (Gray et al.,

1976).

60

The thermal and transport properties of flue gas can be cal-
culated if the composition of the gas is known. The properties are

determined as follows:
1. specific heat (Holman, 1980):

cm = m1C1+ mzc2 + N303 + . . . + mncn (2.43)

2. viscosity (Brokaw, 1958):

 

: E “i (2.44)

where pi an uj arc the viscosities of pure components i and j

(N S/mz).

(2.45)

where: x, = the mole fraction of component i.
xj = the mole fraction of component j.
”1 = the molecular weight of component i.

Mj the molecular weight of component 5.

61

3. thermal conductivity (Mason and Saxena, 1958):

n xi ki
kmix : 131 E (2.46)
1;...
j=1 x3 1)
Jri
¢ij has the same value as that given by equation (2.45)

(4) density:
pi = P/(RiT) (2.47)
1/0 = m1/01 + m2/02 + . . . mn/on (2.48)

where P is the total pressure (bars); R1 the gas constant of component
i (J/kg K); p is the density of a mixture (kg/m3); pi is the density

of component i; mi
temperature of the mixture (K).

is the mass fraction of component i; and T is the

5. Prandtl number:

Pr = Cu/k (2.49)

Due to the large variation of temperature and, therefore, the
thermophysical properties, weighted averages of the above properties

are used. The following equation was suggested by Gray et al. (1976):

62

T
e

"ave g éudT/(Te-Tw) (2.50)

W

where Te and Tw are the flame and the furnace wall temperature (K).

Weighted averages for thermal conductivities, specific heats,
densities, and Prandtl numbers are defined in a similar way as the
dynamic viscosity (equation 2.50).

Adams (1979) devel0ped empirical equations for determining the
thermophysical properties of combustion gases, assuming that the

combustion products behave as air at the same temperature:

om = Mg/RuTe (2.51)
M9 = 30.5 - 12.5 xW (2.52)
“m = (182 + .245)~'r10"7 (2.53)

where M9 is the molecular weight of flue gas; Ru is the universal gas
constant (8.2*10'3 atm per kg-mole per K; Xw is the mole fraction of
water (steam) in the flue gas; and Te is the flue gas temperature (K).
The validity of the above equations is uncertain since the
products of combustion do not behave as air.
Empirical heat transfer correlations have been developed by
several researchers and are generally accepted for estimating convec-

tive heat transfer in tubes. plates and other geometrical shapes

(Holman, 1981). The equations are of the form:

63

Nu c PrmRen (2.54)

Nu

th/k (2.55)

The heat transfer is then given by:

Q = Ahc(Te'Tw) (2.56)

The following is the defining equation for the Reynolds number:
Re = UpL/u (2.57)

The magnitude of the Reynolds number determines whether the flow is
laminar (Re < 2,300) or turbulent (Re > 10,000). Transition from
laminar to turbulent flow takes place in the range of Reynolds numbers

between 2,300 and 10,000 (Holman, 1981).

2.2.2.6.2 Radiative Heat Transfer from Flames

The flame and the combustion products absorb and emit thermal
radiation. Both gases and particulate materials present in the flame
contribute to the absorbing/emitting potential of the flame.

The gases present in flames from biomass fuels include 002,
H20, CO, H2, 02. H, 0, 0H, N2, N and NO. The atomic and the homo-
nuclear diatomic molecules (H,0,N and H2, 02) are the only non-
absorbing constituents (Gray et al., 1976). The rest absorb and emit
radiation over a specific wavelength range or band within the thermal
spectrum. Carbon dioxide and water vapour are the most important

absorbing/emitting gases in combustion products. They are the only

64

two gases considered in engineering calculations; for more precise
calculations other gases should be included (Hottel and Sarofim,
1967).

The range of wavelengths over which absorption occurs in
carbon dioxide and water vapor gases have been given by (Gray et al.,
1976) as 2.4-3, 4-4.8, and 12.5 to 16.5 pm for C02. and 2.2-3.3,
4.8-8.5, and 12-25 pm for H20. The absorptivities and emissivities of
combustion gases vary considerably with wavelength. For engineering
calculations, it is assumed that the gases are grey; thus, average
values of absorptivities and emissivities for the whole of the wave-
length range under consideration are used.

The absorptivity and emissivity of CO2 and water vapour in a
furnace depend upon the flame temperature, the mean beam length of the
gas to the furnace walls, the partial pressure of the gases, the total
pressure, the composition and the amount of the other gases in the
combustion products, and the wall temperature.

Adams (1979) suggested the following equation for calculating
the emissivity of CO2 and H20 in wood-fired boilers:

w = 493[0.6¢A (x +x )1'45/1

9 w co e (2.58)

8c
where e is the combined emissivity of carbon dioxide and water

CW

vapor; A is the grate area (m2); Xw and xco are the mole fractions of

9
the water vapor and carbon dioxide, respectively; and Te is the flame

temperature (K).

65

Soot, fly ash, and particulates in the flame emit and absorb
radiant thermal energy. The emissivity and absorptivity of particu-
lates depend upon their size, shape, concentration, composition, and
the temperature. Also, the wavelength of the radiant energy affects
the emissivity/absorptivity of the flue gases. Adams (1979) suggested
the following empirical equation for the emissivity of soot in a

wood fired boiler:

as = 1-exp(-0.12/Ag) (2.59)

Adams (1979) also suggested an equation for calculating the flame

emissivity from the values of gas and particulate emissivities:

e = as + Esw - (es)(esw) (2.60)

Having determined the emissivity of the flame, the heat radiated to
the wall is given by:

0r = Awea(T: - 1:) (2 61)

2.3 Grain Drying

In this study a concentric vortex-cell biomass furnace (CVCF)
is used to provide heat energy for an in-bin counterflow grain dryer
(IBCF). The dryer was experimentally tested by Silva (1980) using
liquid propane gas. The IBCF drying temperature is between 50 and
100°C depending upon the type of grain dried and the intended use of
the grain (Silva, 1980; Kalchik et al., 1979). When feed corn is

66

dried with air at 65 to 100°C, the IBCF dryer is a high-temperature,
high-capacity dryer. High-temperature drying may affect grain quality
in different ways depending upon the initial grain quality and moisture
content, and the drying/cooling rate.

Grain and seeds are highly perishable if poorly handled. If
well harvested, and stored at a low moisture content and low tempera-
ture, grain will retain the original germinability and other desirable
qualities for a long period of time. In the following sections, grain

drying is reviewed.

2.3.1 Importance of Grain Drying

The importance of grain drying has been discussed by Brooker
et al. (1974). Drying facilitates early harvest, thus, reducing field
losses from storm, insect damage, and natural shattering. Field con-
ditions are often better for harvesting early in the season. Early
harvest permits early and timely seedbed preparation for the next
crop; this is particularly important in some tropical areas where two
or more crops are raised in one year.

Grain drying permits farmers to better plan the harvest season,
and to make better use of labor and machinery since harvesting is not
dependent on fluctuations of the grain moisture content in the field.
Finally, the early harvest enables farmers to take advantage of higher
prices early in the harvest season.

The most important advantage of grain drying is that it permits
long-time storage without deterioration in quality. By removing excess

moisture from the grain, the possibility of natural heating of the

67

grain due to molding and respiration is reduced. Thus, grain via-
bility is maintained during storage. Dried grain (at moisture content
below 13%) is less prone to insect, mites, and fungal damage than wet

grain.

2.3.2 Corn Quality as Affected by Drying Methods

The desirable corn properties are dependent upon the intended
use of the corn. In the US corn is mainly used as animal feed, with
smaller usage as human food, seed and industrial starch manufacturing.
Corn quality is dependent upon several factors: (1) the variety
characteristics. (2) the environmental conditions during growth,

(3) the time and the harvesting procedure, (4) the drying method, and
(5) the storage practice (Brooker et al., 1974). During the drying
process, corn quality is affected by the grain temperature and the
drying rate.

Corn in the US is officially graded for quality under the Grain
Standards Act. The grades and grade requirements are listed in Table
2.5. There are other properties which are of importance to specific
corn users that are excluded from the standard, such as nutritional

value, millability, viability, and susceptibility to breakage.

2.3.2.1 Effect of Drying on Nutritional Feed Value

The most important quality factor of corn for animal feed is
the nutritional value. The effect of drying temperature on the nutri-
tional value of corn for animal feed has received considerable research

attention. Hathaway et al. (1952) found that corn dried at temperature

68

Table 2.5.--Numerical grades and sample grade requirements for

 

 

 

 

 

US corn.
Maximum Limits
Minimum Damaged kernels, %
Grade testweight Moisture BCFM
LB/bu % % heat
damaged Total
1 56 14.0 2.0 0.1 3.0
2 54 15.5 3.0 0.2 5.0
3 52 17.5 4.0 0.5 7.0
4 49 20.0 5.0 1.0 10.0
5 46 23.0 7.0 3.0 15.0

 

 

 

 

 

 

 

Sample grade shall be corn which does not meet the requirements for
any of the grades numbers 1 to 5; or which contains stones; or which
is musty, sour, or burned; or which has any commercially objectionable
foreign odor; or which is otherwise of distinctly low quality.

Source: Brooker et al. (1974).

above 60°C significantly decreased its energy content and palata-
bility. Sullivan et al. (1975) reported that heat has a definite
effect on the nutritional value of corn; also, that the decrease in
commercial quality due to drying at high temperatures may not necessarily
result in a decreased value of corn as animal feed.

Jensen et al. (1960) reported that drying temperatures of 60°C,
82.2°C, and 104°C, have no deleterious effect on the nutritive value
of corn for swine as measured by growth rate and feed use. Jensen
(1978) showed that roasting corn at 15% and 23% moisture at 127°C
and 150°C reduced the availability of lysine; he found that niacin is

unaffected by roasting temperature, but pyridoxine (vitamin 86)

69

availability is significantly reduced in 14% moisture corn when it
is heated at 160°C.

From the above review it appears that corn which reaches
temperatures above 60°C undergoes some minor nutritional changes.
Nutritionists do not agree on the effects of drying temperature on the
feed value of corn (Brooker et al., 1974). It is generally recognized
that physical and chemical properties such as consistency, energy
content. palatability, hardness, color, moisture, vitamins. protein
and amino acid profile are affected by high drying temperature

(Williamson, 1975).

2.3.2.2 Effect of Drying on Corn Milling Quality

Farmers and elevator operators who dry corn often consider
only its feed value. Corn millers are concerned about the increasing
volume of artificially dried corn being marketed (Freeman, 1978;
Rutledge, 1978).

High starch yield (millability), maximum yield of selected
fractions and prime products mix, and low fat content are the most
important desirable characteristics of corn for milling. Brekke
et al. (1973) compared the corn milling response of in-bin natural
air drying with artificial drying in a small experimental fluidized
dryer at air temperatures from 32 to 143°C (maximum corn temperatures
were 32 to 104°C, respectively). The yield of total grits recovered
by sieving, aspiration, and flotation decreased with increasing drying
air temperature; also, the fat content of the grits increased with

increasing air temperatures. Prime products recovered by rolling and

70

grading followed similar patterns; however, sometimes yields and fat
contents were less satisfactory. The results also showed that the
cold paste viscosity of selected products increased if corn dried at
elevated temperatures. The corn dried with natural air had the best
dry-milling quality; drying at 60°C yielded corn of acceptable drying-
milling quality except for a high percentage of stress cracks.

Freeman (1978) discussed the quality factors affecting the
value of corn for wet milling. Drying at high temperatures causes
"case hardening" of proteins. Case-hardened protein affects the
millability by impairing separation and purification of starch. The
result is starch with a high protein content and reduced viscosity.
The drying temperature, drying rate, and the initial corn moisture
content determine the degree of case hardening. High drying tempera-
tures may also destroy some amino acids, especially lysine.

Watson and Mirata (1962) concluded that since kernel viability
is evidently more easily altered by drying conditions than the other
properties examined, corn dried to preserve viability should invari-
ably be suited for starch manufacture. The drying temperature should

not exceed 71°C.

2.3.2.3 Drying Corn for Seed

Generally the techniques used to dry seeds do not differ
greatly from those used to dry grain for other purposes. However,
extra dryer control and management must be practiced in order to
ensure a high degree of germination (Copeland, 1976). The drying air

temperature, the drying rate, and the initial moisture content are

71

the most critical factors affecting the germinating quality. Copeland
stated that the product temperature limit varies with the type of
seed, but should not exceed 38°C; the highest safe temperature also
depends on the initial moisture content. Ulileman and Ullstrup (cited
in Hukill, 1954) showed that seed corn can be dried safely at 49°C, if
the moisture content is less than 25%; for moisture above 25%, 38°C,
is the upper limit.

An excessive drying rate may cause stress cracks. Over-dried
seeds are susceptible to mechanical damage, which is detrimental to

seed quality.

2.3.2.4 The Effect of Drying on Commercial Grade

The effects of artificial corn drying on its composition,
nutritional value, viability as seed, and industrial processing have
been discussed in the previous sections of this review. These factors
are not included in the determination of commercial grade. As shown
in Table 2.5, the only factors considered in the grain standard code
for corn are: (1) the testweight, (2) the moisture content, (3) the
broken corn and heat damaged corn, and (4) the presence of foreign
materials. Artificial drying has a direct effect on the testweight,
the moisture content, and on the percentage of heat damaged corn, and
has an indirect effect on broken corn. These factors will be reviewed

in the following sections.

72

2.3.2.5 Testweight

The corn testweight is its bulk density and is influenced by
grain shape, grain surface texture, moisture content, type and amount
of impurities, size and uniformity, temperature, and other factors
that affect the packing characteristics. According to Freeman (1978),
testweight may indirectly indicate the wet milling quality of corn.
High-temperature drying may reduce the extent of kernal shrinkage
resulting from moisture removal and, hence, result in low testweight.

Hill (1975) studied the effect of drying temperature on test-
weight of shelled corn. According to his observations, the testweight
increases during drying. The increase is due to shrinkage with a loss
of moisture and decrease of the coefficient of friction on the surface,
thus permitting closer packing of the kernels. The testweight increase
is less at higher drying temperatures, possibly due to case hardening.
The testweight reaches a maximum and declines with further drying. The
maximum testweight is reached at 14 to 16% moisture (Brooker et al.,
1974).

The amount of testweight increase with drying depends upon:
(1) the degree of kernal damage, (2) the initial moisture content,
(3) the temperature reached by the grain during the drying process,
(4) the final moisture content, and (5) the grain variety. Early
harvested, high moisture grain is not exposed to much weathering and
shows a higher testweight after drying than the same grain harvested

later at a lower moisture content (Brooker et al., 1974).

73

A higher testweight corn offers some saving in storage since

less storage volume is required to store the same amount of dry matter.

2.3.2.6 Stress Cracks and Broken Corn

Although drying per se does not directly affect the number of
broken kernals, it is well known that grain is physically and physio—
logically damaged when dried at excessively high temperatures. The
degree of damage depends upon the maximum temperature reached by the
grain and the length of the period during which the high temperature
is sustained. The drying and cooling processes directly affect the
degree of stress cracking, and thus determine the susceptibility of
corn to breakage during subsequent handling.

Thompson and Foster (1963) defined stress cracks as the
fissures in the endosperm, or starch inside the kernel, in which the
seed coat is not ruptured. Ross and White (1972) studied the effect of
overdrying in stress cracking in white corn. Their results show a
general decrease in stress cracking as the white corn dried to lower
moisture content, and as drying started at lower moisture contents.
These phenomena are difficult to explain. There may be some physical
and chemical changes such as gelatinization during over-drying which
make the grain kernel more resistant to cracking during the cooling
period. Generally. stress cracking decreases with decreasing drying
air temperature. Slow cooling of both the white and yellow corn after
drying results in a dramatic reduction in the number of checked kernels,
particularly after subjecting the corn at drying air temperatures above

71°C.

74

Gustafson et al. (1978) concluded that there is no significant
increase in breakage susceptibility when corn is dried to 18% moisture
(or above) in a high-temperature dryer; the product of the heating time
and the change of moisture content (under 18%) appears to be the best
predictor of the change in breakage.

Freeman (1973) indicated that broken kernels too large to be
removed by screening for wet milling may release starch granules during
steeping. Free starch in the steeping water causes fouling of evapora-

tor surfaces during steep water concentration.

2.3.2.7 Corn Color

Artificial drying affects other grain characteristics including
color. Ross and White (1972) concluded that darkening and yellowing of
white corn was apparent when it was dried with air at 88°C and 104°C,
and when drying started at an initial moisture content above 25%.
Discoloration was only slight in samples dried at 71°C from any initial
moisture content, and for samples dried from 25% initial moisture con-

tent (mc) to 14% final me with drying air at 104°C.

2.3.3 Drying Systems

There are three basic methods of grain drying: (1) high-
temperature drying, (2) low-temperature drying, and (3) combination
drying. In the US, high-temperature drying has been the primary tech-
nique for more than 25 years. This method is fast, but has a high
fossil-fuel requirement, and can result in a low grain quality. Low-

temperature grain drying (energy may be obtained from electricity,

75

liquid propane, solar energy, or any other heat source) is an energy
efficient process and often results in high-quality grain, if properly
managed. Mold spoilage risk is the main problem encountered in warm
and humid areas. Natural drying is a low-temperature drying method

and takes place when grain is either left standing (or stacked) in the
field, or harvested and dried in a crib. The latter method is often
practiced in the Third World and is the most risky since it exposes
grain to the weather, insects, rodents, birds, and other destructive
elements.

Combination drying processes for drying shelled corn started in
the US in the late 1970's (Brooker et al., 1978). In these processes
high-speed batch or continuous flow drying is combined with low heat
or natural air in-bin drying. The high-speed, high-temperature dryers
dry the corn to a moisture range of 18-23%. The corn is then trans-
ferred to storage where it is slowly dried to a safe storage moisture
content. Combination drying offers a number of advantages, including:
(1) increased output, (2) increased fuel efficiency, (3) improved
product quality (compared to corn dried by high-speed, high-temperature
processes). Brooker et al. (1978) subdivided the on-the-farm high- and
low-temperature drying methods into the following categories: (1) high-
speed, high-temperature batch and continuous dryers; (2) continuous
in-bin drying systems; (3) batch-in-bin drying systems with and without
stirring; (4) low-heat and no-heat in-bin drying systems with and with-
out stirring; and (5) combination systems, in which high-speed batch or
continuous flow systems are combined with low-heat in-bin drying sys-

tems.

76

2.3.3.1 Column Batch Dryers

Column batch dryers are stationary bed dryers, in which the air
moves across a stationary grain column (see Figure 2.7). The dryer is
often portable so that it can be moved from location to location when
not filled with grain. According to Brooker et al. (1974), column
batch dryers have the following characteristics: (1) column thickness
is usually from 30 to 46 cm; (2) column batch dryers operate at high air-
flow rates (over 40 m3/min/m3); (3) drying air temperatures vary from
82 to 116°C; (4) due to the high flow rate coupled with a narrow
column, the moisture gradient across the column is less than with batch
in-bin systems; and (5) drying is completed in a 1 to 3 hour period,
depending on the initial grain moisture content and the need for cool-
ing.

Column dryers are popular because of theirsimple construction
and operation, and because the initial cost is generally lower than of
continuous flow types (Sutherland, 1975). They are suitable for mod-
erate grain volumes (250-650 tons annually) with high initial moisture
contents. Because the dryer has no storage function, it requires well
planned and coordinated handling and storage systems.

The fuel consumption and, therefore, the Operating costs depend
upon the moisture removal range. The fuel efficiency decreases with
decreasing moisture removal range. It requires (at 30% initial mc)
about 5,800 kJ/kg of water removed when the final moisture is 25%, and
6,970 to 8,140 kJ/kg of water when the final moisture content is 15%.

77

Jg" GRAIN HOPPER

 

 
 
 
 
 
 
 

LEVELING AUGER
GRAIN COLUMN

PERFORATED
SCREENS

\

  

HEAT AND COOL

PLENUM

 

 

 

 

 

 

\ /

DISCHARGE AUGER

Figure 2.7. Batch Column Dryer (from Brooker et al., 1974).

78

The thicker the grain column is in a column dryer, the lower is
the specific fuel consumption. However, thicker grain columns result
in an increase in the moisture content gradient across the column.

Kirk (1959) investigated column thicknesses of 10.2, 20.3, 30.5,
and 40.6 cm and concluded that: (1) the 20.3, 30.5, and 40.6 cm
columns are similar in drying air requirements; (2) the operating costs
do not significantly increase with an increase of static pressure of
up to 0.5 kPa for a grain column thickness of 20.3 to 40.6 cm; (3) in
the static pressure range of 0.06 to 0.5 kPa, the drying capacity
increases linearly with static pressure; and (4) there is no signifi-
cant difference in drying capacity with drying column thickness between
20.3 and 40.6 cm.

The moisture and the temperature gradients across the drying
column in a column batch dryer along with the dryer operating costs
can be reduced either by decreasing the drying airflow rate at a con-
stant air temperature, or by decreasing the drying air temperature at

a constant airflow rate (Morey et al., 1976).

2.3.3.2 In-Bin Dryers

In-bin drying systems dry and cool the grain in a bin designed
either as a batch in-bin dryer or as a drying-storage bin. In the
latter case the grain is left in the same bin for storage (see Figure
2.8).

In-bin drying may be categorized in different ways. Brooker
et al. (1974) categorized in-bin drying as (1) full-bin drying,
(2) layer-drying, and (3) batch in-bin drying.

79

 

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80

In full-bin drying a single batch of grain (up to 5 m) is
dried at relatively low airflow. To avoid grain deterioration during
drying, the initial moisture content of the corn must not be above
21-23% (Brooker et al., 1974).

In the layer-drying process an initial shallow layer of grain
is placed in the bin and drying is started. As the drying zone moves
upwards, other layers of grain are added periodically, so that a layer
of wet grain remains ahead of the drying zone. Eventually the bin
is filled; drying continues until the entire grain mass is dried.

This process allows the initial layers of grain to be relatively wet;
they receive more drying air (and therefore dry faster) than the subse-
quent drier grain placed at the top of the bin.

Batch-in-bin drying is a process in which batches of grain are
dried in a bin and transferred to a storage bin. This process utilizes
large quantities of heated air which is forced through relatively
shallow grain beds. Therefore, the drying rate is considerably higher
than in the other in-bin drying systems.

In-bin dryers may be operated with high or low temperatures
(heated air) or with natural air (unheated air). High-temperature
in-bin drying is undesirable in full-bin and layer drying due to
excessive overdrying at the bottom of the bin. However, stirring
devices may be used to solve this problem.

Natural air and low-termperature drying are similar processes
(Bakker-Arkema et al., 1978). The difference is that no heat is added
in case of natural air drying. Low-temperature drying requires raising

the drying air 3 to 6°C above the ambient temperature by either

81

electric heat, solar energy, or other heat sources (Zink et al., 1978).
Liquid propane and electricity are the most common heat sources for
low-temperature drying; both require low capital investment. Liquid
propane gas is usually not used since it requires interval timers to
limit the heating rate.

The airflow rate required for a drying system depends on the
drying system. the harvest date, harvest moisture content, and the
location.

Adding heat, even in small amounts, increases the drying
capacity in a low-temperature drying system (by decreasing the drying
air relative humidity). The temperature increase also results in
faster mold development. To reduce mold growth, the average tempera-
ture in the bin should be below 10°C. With addition of low heat,
the airflow can be limited to 2.4m3/min-m3 for 24-26% mc corn,
1.6m3/min-m3 for 22-24% mc, and 0.8 m3/min-m3 for 20-22% mc corn
(Brooker et al., 1978).

Although low-heat and natural air drying are slow, the quality
of the finished grain is frequently high due to the low application of
heat (Kalchik et al., 1979).

The main disadvantage of in-bin drying is that the grain at
the bottom may be overdried and at the top underdried. Drying is
stopped after the average moisture content in the bin reaches a desired
value. Since unloading does not allow perfect mixing and blending to
obtain a uniform moisture content, the corn may deteriorate during

storage.

82

Several improvements have been incorporated in in-bin drying
to reduce the vertical moisture gradient. This includes grain recir-
culating, in-bin counterflow drying, grain stirring, and drying with
alternating heated and unheated air (Browning et al., 1971).

According to Brooker et al. (1978), the specific fuel con-
sumption of deep bin-in-storage drying is low (3,500 kJ/kg, or lower
in some cases). Therefore, these systems are usually used in combina-

tion with other drying systems.

2.3.3.3 High-Speed Continuous Crossflow Dryers

Crossflow dryers have a wet grain holding bin at the top (see
Figure 2.9). Grain flows by gravity from the top to the bottom through
drying and cooling columns which are 20-45 cm wide. A column thickness
of 28-36 cm is most common. Two fans. one for heating and one for
cooling, are usually used. The drying rate and the final moisture
content are mainly dependent upon the grain velocity, the air tempera-
ture, the airflow rate, and the initial moisture content. Moisture is
usually controlled by regulating the grain flow rate by a metering
auger at the bottom of the dryer, while maintaining a constant air
temperature and airflow rate. The auger speed responds to a tempera-
ture sensor located in the grain column near the lower edge of the
drying section.

The drying characteristics of crossflow dryers are similar to
those of column batch dryers. The grain on the plenum side is over-

dried while that on the exhaust side is underdried.

83

FILLING AUGER

  

 

 

  

FAN
-— AND —— ;:
HEATER

    

HEATED AIR PLENUM

 

 

Gnamcowun

 

 

COOLING AIR PLENUM

 

GRAIN METER
UNLOADING AUGER

 

 

Figure 2.9. Schematic of a Crossflow (Continuous Flow) Grain Dryer
(From Brooker et al., 1974).

84

In some crossflow designs. the cooling air is reused. This
allows recovery of heat from the cooled grain and also reduces the
moisture content gradient across the grain column (Bakker-Arkema
et al., 1980). Some designs incorporate a metering device that
causes the grain on the plenum side to move faster than the grain
on the outside (Rodriguez, 1982). Other designs use a "turnflow
device" to interchange the grain on the plenum side with the grain on
the exhaust side (Fontana, 1983). These improvements are intended to

reduce the moisture content gradient across the grain columns.

2.3.2.4 Concurrentflow Dryers

Concurrent flow drying is a relatively new grain drying tech-
nique (Dalpasquale, 1981; Bakker-Arkema et al., 1983). In a concurrent-
flow dryer, high temperature (BO-300°C) air flows in the same direc-
tion as the grain (see Figure 2.10). The hot air only encounters the
cool and wet grain. Intense heat and mass transfer takes place at the
grain/air inlet. causing rapid evaporative cooling of the air and
heating of the grain. The latter is accompanied by rapid drying of
the grain. The grain temperature remains considerably below the inlet
air temperature. As the grain and the air flow through the drying bed,
an equilibrium is reached (the equilibrium temperature is between the
inlet air temperature and the initial grain temperature). The absolute
and relative humidity of the air progressively increase as the grain
dries.'

The major advantage of concurrentflow drying is that all the

grain kernels are subjected to the same thermal process. Thus, there

Figure 2.10.

mmmonwzb

Wet grain

Drying floor
Exhaust air
Tempering section
Air recirculation
Counterflow cooler
Metering rolls

85

 

 

 

 

H:
I:
J:
L:
M:

 

Schematic of a Concurrent Flow Dryer.

Ambient air

Hot air--First stage
Hot air--Second stage
Ambient air--Cooler
Dry grain

86

is no moisture difference among the kernels. This results in better
grain quality.

Counterflow cooling has usually been combined with concurrent-
flow drying (Bakker-Arkema et al., 1982). Cold air first encounters
the coldest grain, thus limiting thermal stresses.

Several authors have reported the advantages of concurrentflow
drying including: (1) high capacity (in multi-stage concurrentflow
dryers), (2) improved grain quality, (3) greater flexibility for
adaptation to different craps, and (4) high termal effiency (Westelaken
and Bakker-Arkema, 1978; MDhlbauer et al. , 1978; Bakker-Arkema et al. ,
1982, 1983; Dalpasquale. 1982; Fontana. 1983).

2.3.3.5 In-Bin Counterflow Dryers

In-bin counterflow drying is an intermittent process in which
the grain in a bin flows downwards, and the drying air flows in the
opposite direction (see Figure 2.11). The drying process is the same
as for the batch-in-bin dryer. Dried grain is removed from the bottom
of the bin by the means of a tapered sweep auger which removes a thin-
layer of grain from the bottom of the grain bed and delivers it to a
vertical auger; the auger, in turn, discharges the hot dry grain into
a cooling bin. The sweep auger is activated by a temperature sensing
element placed about 46 cm above the false floor. As drying progresses,
the temperature in the region of the sensor increases. When a pre-
selected temperature is reached, the sweep auger is activated; it
makes one complete turn around the bin removing a 7-9 cm thick layer

of dry warm grain. As the auger completes the turn, damp grain

87

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88

moves into the sensor's region and the temperature at that point drops.
The auger is turned off and waits for the next cycle. Thus, the mode
of operation is intermittent. In a countinuous-flow, counterflow
dryer, the sweep auger speed is set so that the grain receives suffi-
cient residence time in the dryer such that the grain is dried to a
desirable moisture content by the time it is discharged. The grain
discharge rate is used to control the final moisture content.

In the counterflow drying process, the warm saturated air
leaving the drying zone passes the cool and relatively wet incoming
grain. Some energy is used to warm the cool grain, and condensation
may occur on the cool grain especially if the bed is deep and the
initial grain temperature is low (Brooker et al., 1974).

Silva (1980), Bakker-Arkema et al. (1980), and Kalchik et al.
(1981) investigated the performance of the in-bin continuous flow
counterflow dryer. The specific energy consumption decreases with
increasing drying temperature (4,390 to 5,110 kJ/kg of water for drying
air temperatures of 93 and 50°C, respectiveIYI; corn quality is not

seriously affected by drying at high temperatures (82-93°C).

2.3.3.6 Combination Drying

In combination drying, wet grain is dried down to 18-24% mc
(wb) in a high-temperature batch or continuous dryer. The grain is
transferred to a bin in which drying is completed by a low-temperature
in-bin drying process (Shove, 1978). Dryeration is a form of combina-
tion drying (Foster, 1964). In the dryeration process, grain is dried

in three phases. The first phase consists of drying the grain down

89

to 2-3 percentage points above the desired final moisture in a high-
temperature dryer. The partially dried grain is transferred to a
tempering bin (second phase) and held hot for 6 to 10 hours. In the
third phase the grain is finally cooled by aeration at low airflow
(.01-.03 m3/min-m3) (Bakker-Arkema et al., 1978; McKenzie et al.,
1972). ’

Gustafson et al. (1976) and Shove and White (1977) showed that
the susceptibility to breakage in the 15-18% mc range is substantially
reduced by eliminating immediate cooling after high-temperature drying.
According to Kalchik et al. (1981), combination drying results in
increased fuel efficiency and increased drying capacity. The average
specific energy consumption for high-temperature drying is 5000 kJ/kg
compared with 3800 kJ/kg for combination drying when corn is dried from

26 to 15.5% mc.

2.3.3.7 Drying Systems Suitable for Biomass Heating

The drying systems discussed previously have been designed for
use with conventional fossil fuels as energy sources for heating the
drying air. A batch dryer has distinct loading, heating (drying),
cooling, and unloading phases during a drying cycle. Thus, the fuel
burner is turned on during the heating phase and is turned off during
the cooling, unloading, and the loading phases. In contrast, continu-
ous flow dryers require a continuous supply of heated air for the
drying sections.

The concentric vortex-cell furnace and other solid fuel burners

require a continuous supply of fuel and combustion air to maintain

90

steady-state conditions. Sudden interruption of either fuel or air
supply results in undesirable transient conditions characterized by
the emission of products of incomplete combustion. Thus, a biomass
furnace cannot be used for heating drying-air for a batch dryer without
incorporation of a heated air diversion system to divert the heated
air to some other use during the loading, cooling, and unloading phases.
Unlike the batch dryers, continuous flow dryers are easily convertible

to the use of biomass heated air.

2.4 Grain Drying Theory and Simulation

Much research has been conducted to study the processes by
which water is removed from biological materials. The drying process
consists of simultaneous heat and mass (moisture) transfer. Heat is
required to provide the energy for evaporating moisture from the drying
grain. The moisture is removed from the surface of the grain by an
external drying medium, usually air.

Bakker-Arkema et al. (1974) listed six possible modes of mois—
ture removal from cereal grains: (1) surface forces (capillary flow);
(2) liquid movement due to moisture concentration differences (liquid
diffusion); (3) liquid movement due to diffusion of moisture on the
pore surfaces (surface diffusion); (4) vapor movement due to moisture
concentration differences (vapor diffusion); (5) vapor movement due
to temperature differences (thermal diffusion); and (6) water and vapor
movement due to total pressure differences (hydrodynamic flow). The

exact manner in which water leaves the grain is dependent upon drying

91

air temperature, air velocity, moisture concentration, and product
type and condition.

Two drying rate periods have been identified: (1) the constant-
rate period during the initial drying of extremely moist grain, and
(2) the falling-rate period. Cereal grains dry during the falling-
rate period (Bakshi and Singh, 1979). The drying rate decreases con-
tinuously as the moisture content decreases during the course of
drying. The thermal and the physical properties of the grain which
are moisture content dependent also change as drying progresses.
Thus, the prediction of the drying rate during the falling-rate
period is more complicated than during the constant-rate period.

The basic drying theory consists of a thin-layer drying equa-
tion (which is equivalent to drying a single grain kernel), coupled

to a series of deep-bed drying equations.

2.4.1 Thin-Layer Drying Models

A "Thin-layer" of grain refers to a layer of grain that is
approximately one kernel deep. A thin-layer drying equation predicts
the drying rate of single kernels of grain as a function of the drying
conditions.

Empirical, theoretical, and semi-theoretical thin-layer equa-
tions have been developed for different crops.

Empirical equations are developed by analyzing statistically
the data obtained by drying thin layers of grain. Theoretical
equations are based upon the diffusion theory (Crank, 1979). Thin-

layer drying equations for several cereal grains have been reviewed and

92

published (Bakker-Arkema et al., 1983; Steffe, 1979; Dalpasquale,
I981; Bakshi and Singh, 1979).

First the thin-layer equations for corn will be reviewed.
Thompson et al. (1968) developed an empirical thin-layer drying

equation for corn in the range of 60 to 150°C.

t = Atn (MR) + B (1n(MR)]2 (2.62)
where: A = 0.004880 - 1.86178
B = 427.36 exp (-0.0330)

The most commonly used thin-layer equations are modified versions of
the Thompson equation (Pfost et al., 1976).
Flood et al. (1969) developed an empirical drying equation for

corn in the temperature range of 2 to 22°C:

MR = exp (-kt0'663) (2-54l
where: MR = (M-Me)/(Mo-Me)
k = exp (~xty)
x = [6.014 + 1.45*10'4(rh)2]'5
-(1.86 + 32)[0.33.41:10'3 + 3111043040215
y = 0.125 - 2.1971110‘3 - [1.88 + 32][2.3*10'5(rh) + 5.8*1o’5]

Rugumayo (1979) developed an equation from mathematical diffu-
sion theory similar to the model presented by Chu and Hustrulid (1968)

for low temperature thin-layer drying of corn. The Troeger and Hukill

93

(1970). the Muh (1974), and the Misra (1978) equations have been
compared by Rugumayo (1979).

In the development of the above equations, it is assumed that
moisture is uniformly distributed within the individual kernels of the
grain. Therefore, the equations cannot be used to predict the moisture
gradient within the kernels. The knowledge of the moisture content gra-
dient within single kernels is needed in order to simulate the tempering
process in grain-drying (Rodriguez, 1982; Fontana. 1983; Sabbah, 1971).

The Crank (1979) spherical diffusion equation can be used to
calculate the moisture gradient in the grain kernels during the dry-

ing process:

0C 33C 20C

'3? ‘ ”('51: * 'r—af) (2'64)
07':

30! 3'14 23”

a " ”Par-1 * a?) ”‘55)

The following are the boundary conditions:

M(v.6) ‘ "111111111
M(o,t) = Me
c(r,o) = cinitial
c(o,t) = C8

where C is moisture concentration (kg/m3) and M is the moisture content

(decimal). The analytical solution of equation (2.64) for the moisture

94

content distribution and the average moisture content of various
regularly shaped bodies can be found 'h1 Crank (1979). For a sphere
the average moisture content is (Sabbah, 1971):

_ 1 "2112 2
MR - '57 exp - [—17— X ] (2.66)

:3401
"MB

n 1

where: X =%-(Dt)2

: The diffusion coefficient (0) is a function of the moisture
content, the grain temperature, and the type of grain. Sabbah (1971)

using a spherical diffusion equation found for corn:

0m = 5.196*10‘3M[exp(-2673.33/8)] (2.67)

A diffusion constant equation should be used with the equation for
which it was developed. The Chu and Hustrulid (1968) equation is a

mass concentration-based diffusivity and is used to solve the equation:

M - M 2 -
H* _ fi =.%gxp[ n/(R Dc)l (2,53)
. e
where: M* = 1'0655Mo - 0.0108

R

radius of a sphere, m

1.51exp {(0.0451a - 5.49)Me - 2513/Ta}

Dc

The grain shape is not assumed spherical; the equivalent radius is

defined as:

95

R - 3V/s

kernel volume, m3

where: V

kernel surface area, m2

V)
II

The temperature and the moisture content Ta and Me’ respectively, are
the drying air temperature and the corresponding equilibrium moisture

content.

2.4.2 Equilibrium Moisture Content Models

The thin-layer equations discussed above require the grain
equilibrium moisture content at the drying air conditions (temperature
and relative humidity). The following are some of the equations used

for predicting the equilibrium moisture contents of corn:

Thompson et al. (1968):

Me = [£n(a - rh/{0.382(1.8T + 82}]'5 (2.69)

Kalchick et al. (1979):

H

0.027“
°°°°2(R”) /1n1 0<RH>52 (2.70)

Me

0.3968xp(.5*RH)/£hT 52<RH>99.99 (2.71)

where Me is the equilibrium moisture content (decimal); RH and rh are
the percent and the decimal air relative humidity, respectively; and
T is the air temperature (°C). Other equilibrium moisture content

equations can be found in Bakker-Arkema et al. (1983). The equilibrium

96

moisture content models should be used with the thin-layer equations

for which they were validated.

2.4.3 Deep-bed Drying Models

For the research presented in this dissertation an in-bin
semicontinuous flow dryer was used to dry corn. The in-bin drying
process can be modeled using a deep-bed simulation model. Bakker-
Arkema et al. (1974) developed a series of deep bed drying simulation
models. The models are subject to the following assumptions:

1. there is no shrinkage of the grain bed during the

drying process;

2. the temperature and the moisture content gradients

within an individual particle are negligible;

3. particle to particle thermal conductivity is negligible;

4. the airflow is plug type (no wall effects);

5. there is no heat loss through the dryer (or bin)

walls and the walls have negligible heat capacity;

6. the heat capacity of the moist air and of the grain

is constant during short-time periods;

7. accurate thin-layer, moisture equilibrium isotherms

and latent heat of vaporization equations are known.

The models (Bakker-Arkema et al., 1974; Rugumayo, 1979) con-
sider the basic laws of mass and heat transfer in the analysis of the
drying process. Thin-layer equations developed by Rugumayo (1979),
Thompson et al. (1968), and Misra (1978), are used to predict the

drying rate of corn. Drying processes of other crops can be modeled if

97

the appropriate auxiliary equations are included. The models solve for
the grain and air temperature, the air absolute humidity and relative
humidity, and the grain moisture content as a function of time and
position in the drying or cooling bed.

The accuracy of the results depends upon the accuracy of the
drying rate equations, and theequations used to predict the thermal

and the physical pr0perties of a given product.

2.4.3.1 Stationary Deep-bed Model

Brooker et al. (1974) analyzed stationary bed drying by making
energy and mass balances on a differential volume located at an
arbitrary position in the stationary bed. The following set of four

differential equations was obtained:

 

 

 

ST -h a (

_ = 1-9) (2.72)
82 Ga ca + Ga CVH

h + c (T-e)
32': h+a *M (T-G) ' f C +v 1M5 .90 (2’73)
”9 Cp 09 cw 0P P pp cw ‘ a 32

3H .. - 32.311

.52 _ 93 at (2.74)
%¥-= an appropriate thin-layer drying equation (2.75)

The boundary and initial conditions are:

a. T(0.t) = Tinlet

b' e(x,o) = einitial

C.

d.

H(o.t) = Hinlet

M(x,o) = Minitial.

98

CHAPTER 3

FURNACE DESIGN, CONSTRUCTION, AND OPERATION

3.1 Design and Construction

The concentric vortex-cell biomass furnace (CVCF) is a modifi-
cation of a furnace developed at Iowa State University and described
by Claar et al. (1981). The furnace consists of two concentric steel
cylinders (see Figure 3.1). The outer cylinder is 1.22 m in diameter
and 3.66 m high and is formed from 11 gage (GA) mild steel. The
inner cylinder is 96.5 cm in diameter and 3.35 m high and made of 3 GA
mild steel. The top 2.75 m of the inner cylinder is lined with 11.4 cm
thick firebricks. The firebricks are supported by a round steel
table with a hole in the middle. The outside and the inner diameters
are 96 and 66 cm, respectively. The height of the table is 56 cm.
The cylinders are covered with lids through which a chimney, 61 cm
high and 36 cm in diameter, made of 3 GA mild steel is inserted. The
inner lid is insulated (on the inside) with 2.5 cm thick blanket
insulation with a thermal conductivity of 3.444 W/m-°C at 800°C. A
16 GA eductor is horizontally connected at the top of the chimney. The
eductor consists of a 147 cm long diffuser with a throat diameter of
36 cm and a diffuser tail end diameter of 66 cm, and an 8-cm nozzle
(see Figure 3.4). The increase in the air velocity at the nozzle tip

and the expansion of the flue gases in the diffuser result in a

99

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101

0.25-0.75 kPa pressure reduction at the top of the chimney inside the
combustion chamber. The pressure reduction depends upon the air
velocity at the nozzle tip.

A 76 cm diameter sloping (12 degree) grate with a cross

2 is located 61 cm above the bottom of the

sectional area of 0.43 m
inner cylinder to support the fuel pile.

An opening in the cylinder below the grate provides access to
remove the undergrate ashes. The opening is controlled by means of two
doors, the outside is hinged on the outer cylinder and the inside is
bolted to the inner cylinder. A 15 cm diameter pipe is located in a
hole cut on the inside door. A disc slides inside the pipe to regulate
the undergrate airflow rate. The disc is adjusted manually by means of
an axial bolt welded on the disc and rotating through a nut welded on
the outside door (the disc-pipe mechanism acts as a gatevalve for
regulating the undergrate airflow). Both doors can be opened to remove
the ash below the grate.

A 25 cm2 square opening_is cut 90 cm above the furnace base in"
both cylinders to provide access to the combustion chamber for lighting
and observation of the fire. The opening is covered with firebricks on
the inside and a mild steel door on the outside.

The inner cylinder (lined with firebricks) has two rows of
tuyeres. three tuyeres per row, positioned at an angle of 18 degrees
with respect to the cylinder wall. The tuyeres are fabricated from

5 cm nominal diameter, schedule 40, seamed stainless-steel pipe. Each

tuyere has a cross-sectional flow area of 22 cm2.

102

The fuel is placed onto the sloping grate by an automatic auger
feeder system. The system consists of an electric motor-driven feed
wagon, a feed hopper and a stoker-auger driven by a 0.37 kw (1/2 Hp)
motor. The rotational speed of the stoker-auger is controlled by a
speed reduction system consisting of 4 to 1 variable speed belt drive,
a 200 to 1 double reduction gearbox, and finally a 4 to 1 chain-drive
reducer. The minimum rpm is 0.5; the maximum 2.

The combustion and the eductor air streams are supplied by a
0.74 kw (1Hp), 36 cm axial fan. The fan forces air into a 36x43x43
cm plenum from which the combustion and the eductor air streams are
supplied.

The combustion air is forced into the furnace plenum (between
the two cylinders) through a 20 cm diameter duct. The air is preheated
to 93-205°C in the furnace plenum by convective and radiant heat
transfer. The undergrate primary combustion air is introduced through
the adjustable gatevalve previously described. Air for secondary
combustion is introduced via the furnace plenum into the secondary
combustion zone through the tuyeres. The combustion airflow is regu-
lated by a butterfly valve.

The air for the venturi eductor nozzle is supplied by a 10 cm
diameter duct connected to the fan plenum. The 10 cm duct is reduced
to 8 tun diameter at the nozzle exit. The eductor airflow depends
upon the combustion airflow rate. If the butterfly valve (regulating
the combustion airflow) is fully open, less air is available for the

eductor, and vice versa.

103

A 1.8 m diameter 24 GA sheet-metal shroud surrounds the top
1.3 m section of the furnace; it extends 1.2 m above the furnace to
improve heat recovery and to facilitate coupling of the furnace to

the in-bin counterflow drying bin.

3.2 Combustion Principle of the Concentric Vortex-Cell Furnace

In the concentric-vortex furnace, combustion air is preheated
between the two steel cylinders. The primary air enters the fuel pile
at the bottom of the grate where gasification and pyrolysis of the fuel
takes place. The producer gases travel upwards into the secondary
combustion zone of the furnace. Secondary air from the furnace plenum
enters into the upper secondary combustion zone through the tuyeres.
The tuyeres cause a vortex and mixing of the flue gases. Unburned
material or fly ash is centrifugally separated from the flame spiral.
The particulates fall onto the top of the burning pile and are reignited.
There are no tuyeres located directly at the top of the burning pile.
The high velocity air from such tuyeres would agitate the pile and
cause excessive unburned particles to accelerate upwards and escape
through the chimney.

The vacuum developed at the tap of the chimney by the eductor
ensures that smoke and other volatile gases do not escape through the
feed auger and other leaks. Thus, the vacuum prevents backfiring
through the feed system and accelerates the flue gases from the combus-
tion chamber into the chimney.

The exhaust flue gases are blended with the ambient air in a

91.4 x 91.4 cm duct that connects the furnace shroud to the dryer fan

104

intake. The temperature of the blended air is between loo-150°C
depending on the desired drying temperature.

The drying system consists of an in-bin counterflow unit built
into a 5.50 m diameter, 3.70 m high bin. It is equipped with a 9 kw
(12.5 Hp) centrifugal fan. The drying system is described in section

2.3.3.5.

3.3 Furnace Operation

The furnace is started by a propane torch and requires about 30
minutes before the desired operating temperature is reached. During
the start-up period, a door in the heat duct between the furnace and the
dryer fan intake is opened to allow the products of incomplete combus-
tion to be discharged into the atmosphere. The dryer fan is turned
on when the exit temperature of the furnace flue gas is about 540°C.
The flue gases are blended directly with ambient air to obtain the
desired drying air temperature. The fuel feed rate is regulated to
maintain the proper drying temperature. The combustion air is kept at
120-170% of the stoichiometric air to avoid smoke and soot in the
exhaust gas and to keep the exit temperature of the flue gas below
700°C. Metal parts in contact with the flue gas experience accelerated

damage when the exhaust gas temperature exceeds 1095°C.

3.4 Concentric Vortex-Cell Furnace Design Theory
Concentric vortex-cell furnaces have been built by trial and
error (Claar et al., 1981). The theoretical basis for the design of

CVCF and other biomass furnaces is lacking in the literature. In the

105

following section the theoretical considerations are discussed that

will contribute to the design of CVCF.

3.4.1 The Design Heat Output

The heat output of a furnace depends upon the environmental
conditions, the furnace design, the grate area, the combustion chamber
volume, and the efficiency of mixing the combustion air and the
combustible materials in the furnace, and the inherent fuel character-
istics (as defined by the particle size, the moisture content, and
the fuel type). Claar et al. (1981) reported a heat output of 5.3
GJ/h per m2 of grate area, of 1.9 GH/h per m3 of combustion chamber
volume. The heat output of the MSU (Michigan State University) CVCF is
2.5 GJ/h (grate area = 0.43 m2, combustion chamber volume - 1.3 m3);
with 20% moisture content of wood, the required maximum fuel feed rate

is approximately 160 kg/h (assuming a net heating value of 15.63 MJ/kg).

3.4.2 Grate Design

The 1981 model of the MSU CVCF had a solid hearth similar to
the hearth by Claar et al. (1980b). It was impossible to continuously
separate the ash from the burning fuel on the solid hearth. The ash
accumulated on the hearth, smothering the fire. Also, the increasing
amount of the ash raised the bed-depth of the fuel pile, resulting in
large amounts of fly ash in the flue gases. Complete combustion of char
was impossible due to the inability of air to penetrate the mixture of
unburned combustible materials and ash. This resulted in inefficiency

due to unburned fuel and in a high resistance to the fresh fuel flow

106

onto the hearth. To overcome the above limitations, the grate was
redesigned in order to ensure complete char combustion and continuous
separation of the ash from the combustible materials.

Junge (1979) suggested the following criteria for designing

grate systems in spreader -stoker boilers:

1. the system must provide uniform distribution of
the undergrate air;

2. the system must permit the removal of ash with
minimum labor;

3. the system must be able to withstand high tempera-
tures;

4. the system must be able to expand and contract with
thermal cycling without buckling, warping, or opening
large gaps;

5. the system must be designed for minimum maintenance.

To satisfy the above requirements, a slightly sloping grate

3 ash receptacle.

(with 12-degree slope) was placed above a 0.25 m
The grate was made of 16 GA, 4-mesh screen. Firebricks were randomly
spaced on the screen to protect it from high temperatures. The screen
(with 6.35 mm holes) was selected to facilitate combustion of wood
chips and corncobs as well as large grains such as corn and soybeans.
During combustion the undergrate air cools the grate and
provides oxygen required for complete char combustion. Combustible

materials are held on the grate while the ashes fall through the grate.

Combustion of char particles is completed under the grate.

107

The 12-degree slope facilitates uniform fuel distribution on
the grate. Also, a combustion front from the lower end of the grate
(and advancing towards the feed) is established. Thus, three regimes
are established in the furnace: (1) char combustion on and under the
grate, (2) new fuel drying on the grate, and (3) fuel pyrolysing and
burning between the char and the new fuel. The undergrate airflow is
regulated to maintain char combustion (to provide pyrolysis heat)
without fluidizing the fuel bed. Fluidization of the fuel bed results
in excessive fly ash development and possibly undesirable combustion

of fuel in the feed hopper.

3.5 Fuel Feed System

The designing of a fuel feed system consists of determining the
feed auger size, maximum rpm and power requirement, and the feed auger
speed control system.

The screw conveyer (auger) designed consists of a standard 25
cm diameter auger rotating in a semicircular trough. The auger is
driven by an electric motor thrdugh the speed reduction system described
in section 3.1.

The theoretical capacity of the auger was calculated using the

following equation (Henderson and Perry, 1976):

108

m = 1.86 x 10‘5 (D2 - d2)Ppr (3.1)
where: m = capacity, kg/h
D = screw diameter, cm
d = shaft diameter, cm
p = pitch, cm
9b = material bulk density, kg/m3
R = auger speed, rpm

Equation (3.1) assumes that the auger is 100% full of the con-
veyed material. The volume of the auger occupied by the biomass
material depends upon the particle size, flowability, abrasiveness,
and packing characteristics of the fuel. The Link-Belt Company classi-
fied bulk solids (including wood chips) for auger conveyor design
(Perry and Chilton, 1973). Wood chips are described as irregular in
size with a sluggish flowability (angle of repose more than 45
degrees), nonabrasive, very light and fluffy, and interlocked to
resist digging. The average bulk density of wood chips is 160-481
kg/m3. Given these characteristics, wood can occupy only 45% of the
available screw conveyor volume. Thus, a 25 cm (diameter) screw has a
maximum capacity of 0.227 m3/h per rpm using equation (3.1) with a
factor of 0.45. This translates to 91 kg/h per rpm if an average
bulk density of 400 kg/m3 is assumed.

The auger used on the MSU CVCF is designed for a minimum and
maximum speed of 0.54 and 2.0 rpm, respectively. Thus, its minimum
and maximum capacities are 49 and 181 kg/h, respectively. The maximum
rate is slightly higer than the 160 kg/h required to give the maximum

heat output of the furnace (section 3.3.1).

109

The power requirement for screw conveyors is the sum of the
power necessary to drive the screw empty and that necessary to move
the material. The first component is a function of conveyor length,
speed of rotation, and friction in the conveyor bearings. The second
term is a function of the total weight of the conveyed material per
unit time, the conveyed length, and the depth to which the trough is
loaded.

In addition to the above components, extra power is required
to overcome transient overloads due to extra friction on the system.
The extra power is necessary when conveying irregularly sized materials
such as wood chips and corncobs. The following equation was used to
estimate the power required to convey the materials (Henderson and

Perry, 1976):
Pk = M L x 1.535 x 10‘4 (3.2)

where PK is the conveying power, kw; M is the material flow rate kg/h;
and L is the length, m.

According to equation (3.2), the power required to convey
181 kg/h for 1 m is less than 0.03 kw. The power required to drive
the screw empty and to overcome transient resistances is considerable.
A 0.37 kw (1/2 HP) motor was selected.

The fuel feed control system consists of two diaphragm switches
(located at the top and bottom of the fuel hopper)and a magnetic relay
switch (see Figure 3.2). The switches control the motor that drives

the feed wagon discharge system. When the feed h0pper is empty, the

110

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111

contacts in the switches A and B (see Figure 3.2) are closed, and,
therefore, the relay switch K is energized and closes contacts K1, K2,
and K3 to energize the feed wagon motor M. Fuel is discharged into
the fuel hopper. Upon reaching level 1 the fuel pushes against the
diaphragm switch A and opens the contact. The current-flow through
the relay is maintained through switch 8 at the top of the hopper and
contact K1. Upon reaching level 2, the fuel pushes against switch B
and opens its contact. Current-flow to the relay is interrupted and
contacts K1, K2, and K3 are opened. The feed wagon motor stops. As
the stoker-auger continues to move the fuel into the furnace, the fuel
level falls below level 2, and contact 8 closes. Since contacts A and
K remain Open, the relay is not energized. When the fuel level falls
below level 1, the pressure on switch A is released and contact A is
closed. The relay is energized and the cycle is repeated. Thus, the
intermittent action of the feed wagon motor maintains the fuel between
levels 1 and 2 in the fuel hopper.

A vibrator (Model 20N manufactured by Eriez Magnetics) is
attached to the side of the fuel hopper to prevent bridging of the
fuel at the bottom of the hopper.

The stoker-auger speed (and therefore, the fuel feed rate) is
manually set to maintain the desired drying air temperature by adjust-
ing the variable speed belt drive. If the fuel conditions do not
change drastically during the combustion process, the fluctuation in

the drying air temperature is limited to i 5°C.

112

3.5.1 Furnace Air Supply System

The furnace air supply system in the MSU CVCF has three func-
tions: (1) to supply oxygen for primary combustion (undergrate air)
and secondary combustion (overfire air); (2) to provide high velocity
air to the eductor for maintaining a vacuum in the combustion chamber
for preventing backfiring (premature ignition of fuel) in the fuel
hopper, especially during start-up: (3) to provide the vortex motion
for centrifugal separation of fly ash and unburned particles from the
flue gas.

The following equation can be used to estimate the stoichio-

metric combustion air (Perry and Chilton, 1973):

Ma = 1.38(C/12 + H/4 - 02/32 + S/32) (3.3)

where Ma is the combustion air (kg/kg fuel dry matter); and C, H,
02 and S are the percentages of carbon, hydrogen, oxygen, and sulfur,
respectively.

The total amount of combustion air (for both the primary and
the secondary combustion) is determined from equation (3.3) using an
average ultimate analysis of wood, and the maximum fuel feed rate
(160 kg/h) previously determined. The average ultimate analysis of
wood (Table 2.2) is 50.5% carbon, 6.2% hydrogen, 46.1% oxygen, 0.88%
ash, and 0.0% sulfur. Using these values in equation (3.3), the
stoichiometric air required is 6.1 kg/kg of fuel dry matter. This is
equivalent to 4.9 kg of air per kg of wood at 20% moisture. Thus, the
stoichiometric air required in the MSU CVCF is 4.9x160 = 784 kg/h at

113

the maximum fuel (wood) feed rate. Excess air is required for complete
combustion due to the poor mixing of air and combustible materials in
the furnace. Also, wet fuels require more air than dry fuels. For
design purposes, 50% excess air was assumed. This gives the maximum
combustion airflow rate of 1180 kg per hour.

The minimum undergrate air may be determined if the ultimate
analysis of char is known. (Table 2.2 contains the available ultimate
analysis of char.) The stoichiometric air required for complete com-
bustion of char is 10.5 kg/kg (using the average ultimate char analy-
sis) or 2.1 kg/kg fuel dry matter since char constitutes about 20%
of the products of pyrolysis (Admas, 1980). Thus, the undergrate
airflow is approximately 4.2 kg/kg fuel dry matter with 50% excess
air.

The eductor airflow requirement may be computed by applying
the Bernoulli equation for adiabatic, nonviscous flow to the eductor
nozzle. The assumption of adiabatic flow is justified by the fact
that the heat transfer per unit mass is small compared to the kinetic
energy changes. The velocity in the nozzle is too large to allow
enough time for the air to come to thermal equilibrium with the walls
of the short eductor nozzle. The Bernoulli equation for compressible

flow is (assuming no elevation differences):

U2 k(P,/o ) U2 k(P2/o )

;+__1_=—£+———2— (34)

2 ’ k-l 2 k-l '
Where U is the gas velocity (m/s); P is the static pressure (kPa);

p is the density (kg/m3); and k is the ratio of specific heat capacities

114

(k = Cp/Cv = 1.4 for air). The subscripts 1 and 2 denote the cross
section at the nozzle entrance and exits. respectively.
The equation of mass continuity can be used to eliminate one

of the unknown velocities, U1 and U2:

Alulp1 = Azuzp2 (3.5)

Algebraic rearrangement of the above equations along with the isentropic

relationship Pl/pE == Pz/p: results in the following equation:

1/2

u2 = [2kRT1C1/C2] (3.6)
. _ (k-1)/k
where. C1 - (Pa/P1)
- 2 Z/k
c2 - [(42/41) “2”.) - 1101-1)
Since k = 1.4 for air, equation (3.6) can be rewritten as:
_ 1/2
u2 - 37.88[T1C1/C2] (3.7)

Equation (3.7) is suitable for adiabatic isentropic flow
through converging/diverging nozzles. A similar equation for straight

ducts can be developed from equation (3.4):

u2 = [2RT1C1k/(1-k)]1/2 (3.8)
For air, k = 1.4 and R = 287, and
u2 = 44.82[T1(-C1)]1/2 (3.9)

115

Equations (3.6) and (3.9) were used to design the axial fan
for both the eductor nozzle and the combustion airflows. For the
eductor to maintain a sufficient vacuum in the combustion chamber, a
pressure of 0.25 kPa below the atmospheric pressure is required. An
axial fan operating at approximately 0.5 kPa above atmospheric pres-
sure was selected.

The eductor nozzle is made of a 10 cm schedule No. 40 steel
pipe connected to a 7.62 cm tapered reducer. The inside cross-sectional
areas of the pipes are 0.0082 m2 and 0.00477 m2, respectively. At an
operating temperature of 50°C, the nozzle exit velocity and the mass
flow rate are 38 m/s and 715 kg/h. A 14-inch diameter, 1 HP axial flow
fan was available with a capacity of 2,100 kg air per hour at 10°C.
With 715 kg/h flowing through the eductor, approximately 1385 kg/h is
available for combustion (some of the eductor air is used to complete
the combustion of unburned volatiles that may be contained in the flue
gases).

Equation (3.9) was used to predict the airflow rate in the
tuyeres. The tuyeres are constructed of schedule No. 40, 2 inch.
nominal diameter pipes with a flow area of 0.0022 m2. The pressure in
the furnace plenum is between 0.125 and 0.25 kPa above the combustion
chamber pressure (slightly below the atmospheric pressure) depending
upon the undergrate airflow rate, the drying airflow and the combustion
chamber temperature. The maximum furnace plenum temperature is about
100°C. Thus, the airflow velocity and the total airflow in the tuyeres
are approximately 20 m/s and 879 kg/h (the undergrate airflow rate under

these conditions is approximately 500 kg/h).

116

The above analysis is used for estimating the fan size for
supplying the eductor and combustion air when the furnace is operated
without the drying system. During drying the drying fan imposes an
additional vacuum on the furnace resulting in additional airflow

supplied to the furnace.

3.5.1.1 Eductor Design
The eductor consists of the eductor nozzle, a suction chamber,
and a diffuser (see Figure 3.3). The eductor is a jet pump in which a
high velocity fluid (the motive fluid) is forced through the nozzle.
The momentum and the kinetic energy of the motive fluid are used to
entrain and pump a second fluid stream, referred to as the induced
fluid. At least three processes are involved (Davieset al., 1967):
1. acceleration of the induced fluid by the impact of
the particles of the‘motive fluid from the nozzle;
2. entrainment of the induced fluid by viscous friction
at the periphery of the nozzle jet;
3. overexpansion of the jet to a pressure below that of
the induced fluid with consequent flow of the induced
fluid towards the axis of the jet.
The design of an eductor requires consideration of the temperature and
pressure of the motive and the induced fluids as well as the eductor
size and geometry.
The diffuser is the most critical part of an inductor. The
diffuser consists of the diffuser entrance, the diffuser throat, and

the diffuser outlet or tail pipe (see Figure 3.3). According to Kroll

117

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118

(1947), entrance to the diffuser throat should be rounded and
bell-mouthed. The taper angle a should be larger than 20 degrees.

This prevents shock and eddy losses at the inlet. The diffuser throat
is the mixing zone of the eductor. Its length should be 4-14 diameters
[the optimum length should be 7.5 diameters (Kroll, 1947)]- The

throat cross-sec'tionalarea should allow the flow of both the motive

and induced fluids without chocking or back flow.

Several methods for determining the Optimum throat area have
been suggested (Kastner and Spooner, 1950; Keenan et al., 1950;

Davies et al., 1967). Theprocedures are especially suitable for
eductors operating at relatively high nozzle pressures (over 34 kPa
gauge). The diffuser tail pipe should have a divergence of 4 to 10
degrees after the throat for pressure recovery. The length of the
tail pipe should be at least twice but not more than 12 times the
throat diameter (Davies et al., 1967).

The distance of the nozzle exit to the diffuser throat affects
the eductor performance (Kroll,_1947). The maximum distance increases
with increased nozzle supply pressure, but should be 0.5 to 5 times the
diffuser throat diameter. Kroll (1947) recommended that the nozzle
position should be adjustable So that adjustment for best performance
can be made during the eductor operation.

The MSU eductor system is designed to operate at a low nozzle
supply pressure (0.5 to 0.75 kPa gauge). Also, the induced fluid
(the flue gas) is energized by compressed air in the furnace and by the

vacuum induced by the drying fan. Thus, the function of the eductor

119

is to induce a slight vacuum in the furnace chimney and, therefore, to
prevent backfiring in the fuel feed system. This means that the eductor
design is not very critical. However, an eductor which induces a high
vacuum in the furnace chimney is desirable since it allows more com-
bustion air into the furnace and thus increases the furnace output.
Several investigators have developed a theory for constant-
pressure and constant-area mixing of the motive and induced air
streams in eductors (Kastner and Spooner, 1950; Keenan et al., 1950;
Davies et al., 1967). The theory is based upon equal temperatures
of the streams. In the MSU CVCF, the motive air is only a few degrees
above the ambient, the induced flue gas temperature is in the order
of 700°C. Therefore, the motive and the induced fluids cannot be
regarded as the same.
The following procedure is developed to calculate the
theoretical eductor throat diameter for constant area mixing. The

momentum equation is, referring to Figure 3.3:

11(pl-p2) = (Ms + MO)U2-M5Ue - MoUi (3.10)

l/k (3.11)

u. = Mo/[p1111 - Ae11 = {Mo/[00(Al-Ae)]}(PO/P1)
where A is area (m3); P is pressure (kPa); M is mass flow rate (kg/s);
U is velocity (m/s); k is the ratio of specific heat capacities,
previously defined; and p is density (kg/m3). The subscripts 1, 2, e,
i, o, and 5 denote the throat entrance, throat exist, nozzle exit,

induced stream just before mixing, induced stream before mixing, and

120

the motive stream, respectively. The nozzle exit velocity Ue is
determined by equation (3.7). Substituting U1 in equation (3.11) and
solving for P2
M u M; P l/k (M + M )u2
P = P +-—§—g + (-9) - 5 O (3.12)
2 1 A1 p0A1(A1 - Ae) P1

A1

 

 

The throat exit velocity, U2, depends upon the temperature, pressure,
area, and the composition of the mixture at the diffuser throat exit.

Assuming adiabatic mixing:

T2 = (MSCSTS + MoCoTo)/MSCs + MOCO) (3.13)
Also:
n
1/0 = Z (Xi/pi)
2 =
2
Therefore,

u2 = (M5 + Mo)/92A1 (3.14)

Substituting equation (3.14) in (3.12):

2 2
M u M (M + M )
s e 0 1/k s
p = p +--———-+- (P /P ) ’ 3.15
2 1 A1 pbA1(A1-Ae) o 1 Aipz ( )

 

 

The throat exit pressure P2 must be at least equal to the
atmospheric pressure on higher to force the mixture out of the

eductor. Thus,

121

2

 

M u M M +M 2
P ”"1514 i .‘ A (3 -11 7(p0,p1)1/k ' {—szo—L = P
1 1‘ 90 1' 1 e 02A1 a
Rearranging:
3 _ _ 2 _
A1(Pa P1) 11105180e + Afi(Pa P1))
M: k. ‘5: Mer
- ... _ - -._&._.. ‘
A1 00 (Po/P1) p2 MsUeAe .02 o (3.16)

Equation (3.16) can be solved for A2. Equations (3.9) to (3.16) can
be used to find the theoretical optimum nozzle/throat area. The
actual optimum area will vary slightly from the theoretical because

of pressure losses due to eddy currents and nozzle/throat roughness.

3.5.1.2 Tuyere Design

The operating principles of tuyeres in concentric vortex-cell
furnaces has been explained by Claar et al. (1980b). The main function
of tuyeres in a concentric vortex-cell furnace is to direct the secon-
dary combustion air into the combustion chamber in a downward vortex
spiral. The downward vortex generated by the air from the tuyeres
encloses an inner upward vortex consisting of a mixture of gaseous
products of combustion, unburned particles, and fly ash. The solids
are centrifugally separated from the inner flame spiral into the down-
ward vortex airstream. The unburned particles are reignited on top
of the fuel bed.

The concentric vortex action minimizes problems with clinkers,

errosion of the firebricks, and corrosive deposits by condensing the

122

slagging fly ash when burning fuels with high fusible ash content
(e.g., husklage).

The upward spiral of the flame increases the residence time
for the entrained particles and combustion gases in the combustion
chamber. This increases the combustion efficiency and minimizes the
escape of unburned particles and fly ash through the chimney.

The design of tuyeres has been by trial and error. There has
been no basic research effort to study the effect of tuyere design
parameters on the reduction of fly ash from and combustion efficiency
of concentric vortex-cell furnaces. The tuyere design parameters
include the tuyere size and angle (the horizontal and the vertical
angle with respect to the furnace wall), the number and the spacing
of tuyeres and the exit air velocity.

As mentioned in section 3.1, six tuyeres are built into the
MSU CVCF furnace. The tuyeres are in two rows, three tuyeres per row.
The first row is located 1.52 m from the base of the furnace. The
spacing between rows is 0.911n,and the tuyeres are equally spaced
around the furnace at an angle of 18° with respect to the furnace
wall. The tuyeres are fabricated from 5-cm nominal diameter schedule
40 stainless-steel. The maximum exit air velocity from the tuyeres is
estimated to be 20 m/s. This design is satisfactory for the MSU

furnace.

CHAPTER 4

THEORY

4.1 Furnace Performance Simulation

The performance of a biomass furnace is measured in terms of
the efficiency of converting the chemical energy stored in the mass
into thermal energy. The major heat losses from a furnace are the
radiation and the convectional heat losses to the environment and the
incomplete combustion of fuel. Incomplete combustion is manifested
by appearance of smoke in exhaust gases and/or carbonaceous material
in ashes.

Biomass furnaces have different configurations of the combus-
tion chamber, the grate, the exhaust system, and the combustion air
supply system. They also differ in the design of the heat recovery
systems. The sizes and the materials of construction are different.
The radiation and convectional heat losses are affected by the con-
figuration and the size of the furnace, as well as the temperature
differential between the furnace and the environment. Thus, it is
impossible to develop a universal simulation model applicable to all
furnaces. The model of the MSU CVCF system developed in the follow-
ing sections can be modified to simulate the heat losses and the

performance of similar furnaces.

123

124

4.1.1 Preliminary Estimates

The conservation of energy equation (2.41) is used to analyze
heat losses from the combustion chamber. If the work and the kinetic
energy terms are assumed negligible, and if the enthalpy terms in
equation (2.41) are replaced by the right hand side of equation (2.42),
equation (4.1) is obtained:

Q =2M(ho - h

TV)“ + AHR +1:M(hTe - ho)“ (4.1)

where M is the mass flow rate (kg/h); ho is the enthalpy of a compound
at 298 K; hTr is the enthalpy of a reactant at temperature Tr (K);
hTe is the enthalpy of a product at temperature Te (K); AHR is the
heating value at 298 K; Tr is the initial temperature of the reactants
(K); Te is the exit temperature of the products (K); the subscripts r,
p, denote a reactant, and a product; and the subscript i denotes a
compound. Equation 4.1 can be used to predict the adiabatic flame
temperature (or the total heat losses if the final product tempera-
ture is known). In a biomass furnace the reactants are fuel dry
matter, fuel free moisture, and air. The major products are carbon
dioxide, water, nitrogen, and excess oxygen.

The MSU CVCF furnace is analyzed using Figure 4.1.

The energy balance on control volume 1 (CV1) results in

equation (4.2):

z:M(h .-h ).= 2 M(H -h ).
prod Tu Tu 1 reac 298 T3 1
+ AHR + - z M(hT -h298)1. (4.2)

prod u

125

 

 

 

 

 

 

 

1.
-' 8 DILUTED
COMBUSTION 8 FLUEGAS
AND r-'----—’_-—— "r
Bounce-1.; 5 6, '
AIR '
‘ CV1
. ------ .m/
l . . 1 M
l 1
| l : I
' ' ‘ 1 1
13 1
l +<DMBUSTION +1 '
: 1 « ; 1 1
I
1 , 1 1
1 ' ‘ I
‘ 1-- - - .1181! ..... 1 l
L--..-__ ____- ___--1

Figure 4.1.

Schematic Flow Diagram of the MSU Concentric Vortex—

cell

1, 2
3
1, 7
2, 5
8

Furnace.

Ambient air

Preheated combustion air

Diluting air drawn around the furnace
Eductor air stream

Diluted flue pass to dryer

126

The adiabatic temperature T; is computed by equation (4.1) with zero

heat losses:

).- o (4.3)

2: M (hm- hT3)i+ ARR + a 1

2: M (h .- h
reac T

prod 1 29

The reactants in equations (4.2) and (4.3) are the air mass stream
M3 and the fuel. Similar equations are obtained for control volume 2

(CV2):

2: M (hT.-hT )1. = 2 M (h -hT )1. + AHR

prod s e react 29° 2 o
+ Z M (hT -h ). (4.4)
prod 6 293 1
2: M(h -h ).+AH + z M(h .-h ).=0 (4.5)
reac T2 1 R0 prod T5 29° 1

The reactants in equations (4.4) and (4.5) are the air mass stream
M2 and the fuel.

The heat output from the combustion chamber is the sum of
the energy required to preheat the combustion air and the fuel, the
energy required for preheating the diluting air before mixing, and the
heat lost to the environment. The heat input into the furnace is the
product of the fuel dry matter feed rate (kg/h) and the fuel higher
heating value (kJ/kg). Thus, for control volume ;CV1:

q + Msc (Ta-T2) + MIC (T7-T1) - ~73 MiAhi'r +4111, + >3 1414er (4.6)
b

reac prod

127
and for control volume CV2:

Q + M10 (T7-T1) a .- 2 MiAhi + AHR + Z M10111 (4.7)

reac T prod T
2 s

where Ahi = hi - hi

T. T.

J J 298

C = specific heat (kJ/kg°C).

The mass flow rates and the ambient temperature are known in
equations (4.2) to (4.7). The equations are solved by an iterative
procedure that starts with an estimated temperature of the preheated
combustion air T3. Consequently, the adiabatic flame temperature can
be computed from Equation (4.3). The corresponding furnace tempera-
ture Tl+ is obtained from equation (4.2).

The eductor exit temperature is computed by applying mass
and energy balances after mixing the flue gases and the eductor motive

air at point 5. Thus,

T: = ”“2555 I 1:384) <4-81

The temperature of T6 obtained from equation (4.8) is compared
with the solution for T6 in equations (4.4) and (4.5). The iteration
is repeated until a solution is obtained. This procedure results in
approximate values for the combustion air preheated temperature T3.
the furnace combustion chamber temperature T“, and the eductor exit
temperature. The total heat loss cannot be predicted since equations

4.6 and 4.7 cannot be solved.

128

The final combustion chamber temperature, the preheated com-
bustion air temperature, preheated diluting air temperature, the
furnace flue gas temperature before and after dilution, the drying air
temperature and the total heat losses can be computed from a heat

transfer analysis as will be shown in the next section.

4.1.2 Heat Transfer from Combustion Chamber
The heat losses from the combustion chamber occur in the form

of radiative and convective heat transfer.

4.1.2.1 Radiative Heat Transfer
The flame in the combustion chamber looses heat to the chamber
walls by radiation and convection. In the model the radiative heat

transfer is predicted using the following equation:

1, 11
QR = AwEFw-fflf ' Tw) (4'9)

The flame emisSiVity is computed using the Adams (1980) equations for
soot, carbon dioxide, and water vapor (equations 2.58-2.60). Since
the flame is entirely enclosed within the furnace chamber, the view

factor Ff—w is equal to one.

4.1.2.2 Convective Heat Losses

The convective heat transfer from the flame to the combustion

chamber walls is calculated by:

Qe = Awhc(Te - Tw) (4-10)

129

The convective heat transfer coefficient hc is estimated from

the empirical correlation for rough pipes (Holman, 1981):

2
StbPrf ’3 = f/8 (4.11)

Stb = hC/(pCUm) (4.12)

Equations (4.11) and (4.12) are rearranged to solve for the convective

heat transfer:

11C = prUm/B (4 13)

In equation (4.13) f is the friction factor. The friction factor
depends upon the roughness of the surface and the Reynolds number, and
varies between 0.008 and 0.1. In the MSU CVCF furnace the combustion
chamber is lined with bricks and, therefore, it can be regarded as a
rough pipe. The maximum value of 0.1 was selected for the friction

factor of the chamber.

4.1.2.3 Total Heat Transfer from Combustion Chamber

The net energy out of the furnace is the sum of the enthalpy
change of the reactants (fuel and air) from the initial temperature to
the flame temperature, and the net heat transfer from and to the cham-

ber walls. The next heat transfer is given by:

Qw = Aw[(ewa_fo(T; - 1;) + hc(Tf - Tw)] (4.14)

130

4.1.3 Radiative Heat Transfer Between Cylinders

The furnace jacket consists of the annulus region (plenum)
between the inner and the outer cylinders (marked D and E in Figure
4.2). Heat is transferred from the surfaces of the inner cylinder (6)
to the outer surfaces of the cylinder (7) by radiation and convection.
The surfaces (6) and (7) are considered concentric cylinders of infinite

length. The radiation shape factors are (Howell, 1982):

F5_‘7 = 1 (4.15)

F7-5 = rG/r7 (4.16)

F = 1 - F (4.17)
7-7 7.5

 

 

u u
- 0 (T6- T7)
Q6-7 - 1-66 1 1'87 (4°18)
+ +
A666 A6F6_7 A76,

The radiation thermal resistance between surfaces 6 and 7 is:

1 - as + 1 + 1 - e7
MFG”7 A787

 

(4.19)

 

1s - A566
Thus, equation (4.18) can be written as:

o (T“ - T“)
Q6 7 = “R 7 (4.20)

16

 

131

 

 

 

 

 

 

Figure 4.2. Simplified Furnace Diagram. The letters refer to spaces
and the numbers to surfaces.

132

Figure 4.3 represents the radiation, the conduction, and the convection
heat transfer network for the plenum area between the inner and the
outer cylinders, and the furnace combustion chamber. Heat balances on
the network of Figure 4.3 result in the following set of simultaneous

equations:

on surface 7:

T'Eb (%—+%—)-Qc -QC -—=o (1.21)

On surface 6:

 

1.4. _ 11..- a...
R R
18 16

- Q = o (4.22)

C6

0n the furnace wall W:

Ebf Ebw Tf Tw Tw TE
R C R
19 f » 18

The radiative heat transfer between the surfaces 1, 2, 3, and
4 is complex since all the surfaces radiate and absorb heat to and
from each other. Surfaces 1, 2, 3, and 4 are formed by two coaxial
cylinders (of inner and outer cylinders 2 and 4, respectively) and the
two annular ends, 1 and 3. The radiation shape factors between vari-
ous surfaces are calculated using the governing equations for the

relevant shapes (C-86, C—87, and C-89) as presented by Howell (1982).

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The radiative heat transfer among the surfaces 1, 2, 3, and 4 can be
simplified by analyzing the radiation network shown in Figure 4.4.
The radiation resistance R1 to R10 are defined in Figure 4.3. The net-
work in Figure 4.4 is analyzed by the application of Kirchhoff's cur-
rent law which states that the sum of the energy flows into a node is
zero. The Ebi terms in Figure 4.3 represent the product ng, the
radiocity from a node i (the total radiation which leaves a surface
per unit time and per unit area).

By applying the Kirchhoff law to the network in Figure 4.4,

a set of four simultaneous equations is obtained:

ELL-fiZ-fJ“ +J3_J (L+_1_+_L_+_1__)=0 (424)

R R R R 1 R1 R2 R1, R10 '

1 2 10 "1

%Z+J.+i1.i1-, 1 1 1.1)... (425)

R3 R2 R9 R 2 R2 R R R9 .
J J J

ii+—1—+—2-+-3-J —1—+—1—+i-+-l-)=0 (4.26)

R7 R10 R8 R6 11 R6 R7 R8 R10

E J J J

_L 4 4 4- Jqlqlutho Mfl)

R + R. + R9 + R5 J3 (R. R5 R5 R9)

A fifth equation is obtained by writing an energy balance for

surface 4 that includes the convective heat loss from the surface:

____£.1. .. - -
R7 QCu-o QCu-i (4 28)

135

 

 

 

 

 

 

 

 

 

R
1 J1 R2 J2 R3
Ebl Eb2
R“ Re
3 E
b3
Rs 33 Rs J1 R7 b“
R = 1'5 R a 1 R = 1’6 R = 1
1 A c A F A e A F
1 1 1 1'2 2 2 1 1'3
-e
R5 8 1 3 R 8 1 R 3 1 R 8 1
A e 5 A F 7 A 9 A r
3 3 3 3‘4 3 3“» 2 2-1
R =- 1 R I 1
9 A F 1° A F
2 2-3 1 1-1

Figure 4.4. Complete Radiation Network for the System 1, 2, 3, and 4
in Figure 4.2.

136

4.2 Heat Transfer Between the Furnace and the Hood

The surface marked 9 in Figure 4.2 is a reflective surface.
Its emissivity and absorptivity are approximately zero. Thus, the
thermal radiation resistance is high. Therefore, the four-surface
system (3, 4, 8, and 9) reduces to a three-surface system consisting
of surfaces 3, 5, and 8. The radiation network for the reduced sys-

tem and radiation resistances R11 to R1 are presented in Figure 4.5.

S

The Kirchhoff analysis of the network in Figure 4.5 results

in three simultaneous equations:

__S+__?_+_i.-J L+;+-—1—- =0 4.29
R5 R13 R11 5 (R5 R11 R13) ( )
E J J
bs 5 7 1 1 1
__+.__+ J -——-+—+— =0 4.30
12 R11 le 6 (R11 R12 R14) ( )
E J J
be 5 5 1 1 1
_. ....-.) __+—_+— =0 4.31
R1s -§13 R11 7 (R13 R11 R15) ( )

The energy balance on surface 3 (including radiation and convective

heat gains and losses on both sides of the surface) gives:

Ja-Ebs-EbJL-J _Q
R5 R5 C3‘0

 

- Q = 0 (4.32)

Equation (4.33) is obtained from an energy balance on the top plate 5
(including convection from the top of the plate to the ambient and to
the space between plate 5 and 3, and the radiation from the plate to

the sky):

137

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J - E E - E
5 b5 b5 b3 _
R12 Ra ' ”WATS-Tm) - 1151-0545141— 0 (4.33)

 

 

Equations (4.21) to (4.33) consist of a set of thirteen simul-
taneous equations with 13 unknowns. A computer program was written to
solve the set of the equations after estimating initial values of T2,
T7, and the temperature of the preheated combustion air Td as functions
of the combustion chamber temperature. The estimates are based on sta-
tistical analysis of the experimental temperature data. The following

are the equations used for the initial guesses of the temperatures:

T, = 0.691f - 24.10 (4.34)
T, = 0.241f - 139.80 (4.35)
Td = 9.181f - 83.00 (4.36)

Equations 4.1 to 4.8 are used for the initial guesses of the preheated
air temperature, the furnace adiabatic temperature, and the actual
flame temperature.

The convective resistances and heat transfer rates are calcu-
lated using standard procedures for the relevant flow regimes, shapes,
and temperatures as presented in the heat transfer literature (Holman,
1981; Gebhart, 1971). The convective heat transfer coefficient on the
outer surfaces is a function of the ambient temperature. The influence
of wind and precipitation are not included in this analysis.

Ixcomputer program was developed and used to analyze the furnace

performance in terms of the heat losses and heat output at different

139

fuel feed rates, and at different fuel and ambient conditions. The

inputs to the model include:

a.
b.

C.

The fuel type.

The fuel moisture content (% wet basis).

The fuel feed rate (kg/h).

The ambient temperature (°C) and the relative

humidity (percent).

The drying bin plenum pressure (kPa) (the pressure is

employed to calculate the drying airflow rate).

The program solves for the maximum flame temperature and for

the wail temperatures contained in equations (4.21) to (4.33). The

output of the program includes (see Appendix A):

1.

The
a.
b.

C.

temperatures in and around the furnace:

The preheated combustion air temperature.

The furnace combustion chamber (flame) temperature.
The exit flue gas temperature.

The diluted air temperature (immediately after
mixing flue gases with ambient air).

The drying air temperature and absolute humidity.
mass flow rates in the system (kg/h):

The fuel feed rate.

The combustion air.

The drying air.

The ash accumulation (and fuel accumulation if the

combustion air is less than stoichiometric).

140

3. The total thermal energy input (MJ/h).
4. The total heat loss from the system (Md/h).

5. The furnace and the system efficiencies (percent).

4.3 Grain Drying Simulation

The in-bin intermittent counterflow drying process is simulated
using the MSU fixed bed simulation model (Bakker-Arkema et al., 1974),
equations (2.72) to (2.75). The equations are solved by finite dif-
ference techniques. The Troeger and Hukill (1970) thin-layer equa—
tion is used for drying air temperatures less than 71°C, and the
Thompson et al. (1968) equation (2.78) for grain temperature above 71°C.
The DeBoer equations (Bakker-Arkema et al., 1974) are used to esti-
mate the equilibrium moisture content required in the thin-layer equa-
tions.

The intermittent in-bin counterflow dryer is similar to a fixed
bed in-bin dryer. The grain is held in the bin until the bottom layer
(7.5-10 cm deep) has dried to a predetermined moisture content. The
dry grain is then removed and the process continues until the bin is
empty (if grain is not added to the bin). This process is simulated by
an in-bin fixed bed drying model which reduces the bed depth instan-
taneously by the equivalent depth of the grain removed and increases
the airflow rate accordingly. The grain conditions at various levels
above the false floor are maintained and shifted downwards as drying
progresses.

During intermittent counterflow drying, the drying airflow

increases with a decrease in grain depth and decreases each time some

141

grain is added to the bin. To simulate these changes, the character-
istic fan equation, which predicts the airflow delivered by the centri-
fugal dryer fan at a given static pressure, and the resistance of

the grain to airflow should be known. The characteristic fan equation

(4.37) was supplied by the manufacturer (Parkes, 1982):

v = a - 8P° (4.37)
where Y = volumetric airflow rate (m3/min)

P = pressure (kPa)

a = constant (368.81 m3/m1n)

B = 80.87 0 < P < 1.9

8 = 56.23 1.9 5_p ( 2.4

C = 1.17 0 < P < 1.9

c = 1.95 1.9 E P < 2.4

The pressure drop due to the grain (corn) airflow resistance
is predicted by equation (4.38) (Hukill and Shedd, 1955):

_ 7 x 10'3 dTy

P ’ 1n(1 - 0.512y)

 

(4.38)

where y is the airflow rate (m3/min - m2), and D is the corn depth (m).
An additional constant 1 was introduced in the original Hukill and
Shedd equation to account for fines, corn kernel size, parking of the
corn in the bin, and foreign materials. The value of the constant T

is found by solving equation (4.38) using experimental values of pres-
sure (P) and grain depth (D). The value of T for the corn at the

Kalchik Farms was found to vary between 1.3 and 1.75.

142

Equation (4.38) and (4.37) may be combined to a single
equation (4.39):

3

. l/c '
A — - 7310 D72 _._ 4 .39
(Jr—2) m(1—o.512y) O ( )

Thus, for a given grain depth there exists y > 0 which satisfies
equation (4.39). The value of y is found by the secant iterative method
for finding the roots of “an equation (Conte and de Boor, 1980). The
pressure drop is computed by equation (4.38), using the y value that
satisfied equation (4.39).

The furnace thermal analysis model and the in-bin counterflow

grain drying models are included in Appendices C and D, respectively.

CHAPTER 5
EXPERIMENTAL INVESTIGATIONS

5.1 Furnace/Dryer Performance

The MSU concentric vortex-cell furnace (CNCF) was tested in
conjunction with an in-bin counterflow dryer. The fuels used were
wood chips at 14, 19, and 45 percent moisture content, and corncobs
at 10 and 12%. Shelled corn was dried from different initial moisture
contents (22-27%) to approximately 17-18% and dumped hot in a cooling
bin in which it was held for 6 to 12 hours before cooling at a low
airflow rate (0.5-1.0 m3/min-m2). Moisture contents after cooling
varied from 14 to 16%. The airflow in the drying bin varied from 11
to 13 min3/min- m2; the drying air temperature from 65 to 94°C. The
airflow in the cooling bin varied from 0:5 to 1 m3/min-m2; the cooling
air temperature from 0 to 13°C.

During each test the following temperatures were measured and
recorded:

1. The temperature in the furnace plenum and above the

grate (after primary combustion).
2. The temperature between the two sets of tuyeres.
3. The temperature above the top set of tuyeres.

4. The temperature at the chimney exit.

143

144

5. The temperature at the dryer fan-intake after

diluting the flue gases with ambient air.

6. The temperature in the dryer plenum.

In addition, wet and dry bulb temperatures of the exhaust air after
drying were measured along with ambient wet and dry bulb temperatures.

The furnace temperatures were recorded using Type K (chromel-
alumel) thermocouples. All other temperatures were measured using
Type T (copper-constantan) thermocouples. The limits of error are
i 0.75% and i 0.83°C for Types K and T, respectively.

The temperatures were recorded on a Kaye Digistrip II recorder
at one minute intervals during start-up, and every five minutes there-
after.

The furnace airflows were measured using Alnor-Velometer series
6000-P airflow meter (error limit 1 0.5 m/min). The blended drying
airflows were calculated from the characteristic fan curve after
measuring the plenum pressure.

The fuel feed rate was determined by recording the time
required to burn a known weight of fuel. The corn and the fuel mois-
ture contents were determined using the standard oven method (ASAE

1982).

5.2 Chemical Analysis

The corn dried with the CVCF/IBCF system was analyzed for
polycyclic aromatic hydrocarbon (PAH) and heavy metal contamination.
The corn samples for the contamination analyses were taken at the

dryer inlet and during each cycle as the grain was removed from the

145

bottom of the IBCF dryer and discharged into the cooling bin. The
cycle time was approximately one hour. The initial corn depth varied
from 1.3 to 2 m. The residence time of the corn in the dryer varied
from two to ten hours. The corn that was transferred to the cooling
bin in the first cycle had the shortest residence time, the corn
transferred in the last cycle had the longest.

A propane test was conducted in a batch column dryer for com-

parison.

5.2.1 Determination of Corn Contamination with PAH

The contamination of corn with PAHs was analyzed using a high
performance liquid chromatography (HPLC). The chemicals and the
reagents used for the analysis are described below.

Hexane and methylene chloride of analytical reagent grade were
used along with acetonitrile distilled in glass of U.V. grade. Water
was distilled, deionized, purified of organics, and filtered through
a 0.22u pore size membrane. Polynuclear aromatic hydrocarbon reference
standards were obtained from Supelco, Inc. (Bellefonte, PA). Silica

gel (SilicAR cc-7 special) was employed for the column chromatography.

5.2.1.1 Extraction and Purification of Samples

Samples were extracted with methylene chloride in a soxhlet
extractor and purified by silica gel chromatography using procedures
described by Jacko et al. (1982). Shelled corn (20 g) was extracted
with 125 m2 methylene chloride for six hours, with the temperature

adjusted to give a cycle time of six minutes in the soxhlet. The

146

methylene chloride was evaporated on a rotary evaporator, and the
sample redissolved in Smt of hexane. A 1.2 x 28 cm column containing
4 g silica gel was topped with 2 cm of anhydrous Na2504 and equi-
librated with 25 m2 haxane. The corn sample was rinsed into the
column with 5 m2 of hexane and the column eluted with 25 m2 of
hexane which was discarded. Subsequently, the column was extracted
with two 25 m2 fractions of CHZCl2 and hexane (40:60) which were
combined and evaporated to near dryness on a rotary evaporator. The
sample was transferred to a test tube with CHZClZ and evaporated

to dryness in a stream of N2. The samples were redissolved in 2 m2
of acetonitrile and filtered through a 0.4u pore size membrane prior

to analysis by high performance liquid chromatography (HPLC).

5.2.1.2 High Performance Liquid Chromatography

The HPLC analysis was carried out as suggested by Ogan et al.
(1979). The HPLC equipment consisted of a Waters Associates (Waltham,
MA) Model 680 automated gradient controller, a Haters H6000 and M45
solvent delivery system with U6K injector, a Perkin-Elmer LS-4
fluorescence detector and a Hewlett-Packard 3390-A integrator recorder.
The column was a Beckman Ultrasphere 3u-ODS (reverse phase), 4.6 mm
1.0., and 7.5 cm long. The solvents employed were pure acetonitrile
(Solvent A) and acetonitrile/water mixture at a 20:80 ratio (Solvent B).
The chromatography was performed at a flow rate of 1.5 mzlmin. The
samples were eluted and the column re-equilibrated under the following

program SEQUEHCE:

147

0-20 min: 40% to 100% Solvent A

20-30 min: 100% Solvent A

30-35 min: 100% to 20% Solvent A

35-40 min: 20% to 40% Solvent A

40-55 min: 40% Solent A, isocratic

Polycyclic aromatic hydrocarbons were detected at excitation
and emission wavelengths optimal for each compound as determined by
preliminary scans of the reference materials on the LS-4 detector.
The entrance and the exit slit widths were 10 nm each. Table 5.1
contains the wavelengths for the fluorescence detector used to deter-
mine the PAHs, along with the expected retention time and the detec-
tion limit of each compound. Actual elution times varied slightly
from day to day, but were determined by establishing a reference mix

before each unknown.

Table 5.1.--Excitation/emission wavelengths, retention times, and
detection limits of different PAHs on the fluorescence

 

 

 

detector.
Eggggggggn/ Retention Detection

Compound wavelength time limit

(nm) (min) (ppb)
Naphthalene (Nph) 261/328 4.4 1.25
Fluorine (F1) 257/305 8.6 0.12
Anthracene (An) 251/400 10.8 0.06
Fluoranthene (Ft) 235/388 13.6 0.06
Benz [a] anthracene (BaA) 281/386 17.6 0.06
Benzo [a] pyrene (BaP) 305/410 20.0 0.06

 

 

 

 

 

148

5.1.2 Determination of Corn Contamination with Heavy Metals

The concentration of heavy metal on the biomass and the propane
dried corn was determined using inductively coupled argon plasma atomic
emission spectroscopy (ICP-AES).

Intra-analyzed, concentrated nitric acid was employed along
with yttrium of analytical grade. Water was the same as described in
section 5.2.1

One-gram corn samples were wet-ashed with 5 m2 of concentrated
nitric acid in 30 mt containers and kept for 12 hours in a low tem-
perature (75-80°C) oven. The samples were transferred to 25 m2
volumetric bottles spiked with 10 ppm yttrium which acted as an internal
standard. The samples were diluted to a volume of 25 mt with distilled
water.

The plasma spectrometer (ICPaAES) employed for heavy metal
analysis was a Model 955 Atom Comp manufactured by Jarrel-Ash Division
(Naltham, MA). The results were printed directly in ppm. The follow-

ing were the conditions during each analysis:

Coolant (Argon) flowrate 18.0 t/min
Sample carrier flowrate 0.5 £/min
Sample flowrate 1.0 t/min
Nebulizer pressure 1.8 bars
Forward power 1 . 1 kw
Refelected power < 5 N

The detection limits for the heavy metals were the following:
arsenic 3.0 ppm, cadmium 0.3 ppm. chromium 0.6 ppm, mercury 6.0 ppm,
selenium 15.0 ppm, lead 3.0 ppm, and thallium 15.0 ppm.

CHAPTER 6
RESULTS AND DISCUSSION

6.1 Introduction

The primary objectives of this study were to develop and
incorporate a concentric vortex-cell biomass furnace (CVCF) in an
existing on-farm grain drying system, and to test and analyze the
furnace/dryer system in order to determine its economic feasibility
and safety aspects.

In this chapter the experimental and the simulated results
will be compared. The simulation models are used to predict the fuel
requirement and the corn drying rate using different fuels under
different ambient conditions. An economic analysis is made for average
conditions.

A corn contamination analysis is presented to assess the

health hazards associated with biomass fuels.

6.2 Fuel Consumption

Table 6.1 contains the experimental and the simulated fuel
consumption rates for different ambient and drying conditions. The
table also contains the minimum theoretical fuel consumption rate
required in the case of zero heat losses from the furnace/drying

system.

149

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151

The fuel consumption rate depends on the heat required to raise
the drying air from ambient to the drying temperature, on the heat
losses from the system, on the fuel heating value, and on the fuel
moisture content. The fuel consumption varies slightly with the
combustion air/fuel ratio (and therefore with the excess air).

An in-bin counterflow drying system (IBCF) was operated at a
2 m bed depth, an airflow rate of 300 kg/min, a drying temperature of
77°C, and at 12.4°C ambient temperature. This requires a minimum
energy input of 1213 MJ/h (without considering heat losses). The
energy input is equivalent to approximately 78 kg/h of wood chips or
74 kg/h of corncobs with 14.6 and 12.1 percent moisture (wet basis),
respectively, if the average heat values of these fuels are used (20.97
and 18.65 MJ/kg, respectively). The minimum theoretical fuel feed
rate for other tests is similarly computed and is presented in
Table 6.1.

The actual fuel feed rate is the sum of the theoretical minimum
feed rate and the feed rate required to offset the energy losses from
the system. The experimental and the simulated fuel feed rates are
included in Table 6.1.

Table 6.2 contains the experimental and the simulated heat
losses from the furnace/dryer system for the tests and ambient condi-
tions presented in Table 6.1.

The experimental heat losses are based upon the experimental
furnace temperatures (of furnace surfaces and of the diluted air leav-

ing the furnace), the drying air temperature in the drying bin plenum,

152

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153

the ambient temperature, and the drying air mass flow rate (kg/h).
The average experimental heat losses are calculated as follows

(using test No. 1 results for illustration):

Average experimental temperatures (Test 1):

Top of furnace shroud 39°C
Furnace side wall 91°C
Diluted air at furnace exit 99°C
Drying air (bin plenum) 77°C
Ambient 12°C
Sky temperature -7°C

The sky temperature is estimated by an equation presented by Duffie

and Beckman (1980):
_ 1.5
TS - 0.0552Ta (6.1)

where TS and Ta are the sky and the ambient temperatures (K), respec-

tively.

154

1. Furnace radiation loss:

1 (a) From top of shroud (radiation):

4
S)

er(T: - T (6.2)

Qrt

5.66 x 10"8 x 2.6(3124 - 2664)
656 w (2,362 kJ/h)

1 (b) Wall radiation:

4 T4)

0.5A [F e (Tw - a

rw ' w w-g w-g

O
I

4 4
+ Fw_sew_s(Tw - 75)] (6.3)

3.3[(3124 - 2854) + (3124 - 2664)]

1,386 N (4,991 kJ/h)

Total radiation loss - Qrt + Qrw = 7,353 kJ/h

2. Furnace convective heat loss:

From top of shroud (convection):

Tfilm = (39 + 12.4)/2 = 25.7 0 (299K)
- 3 2

Gr - gB(Tw — Ta)L /v (6.4)
= 1.37 x 1010

Ra - PrGr = 0.708 x 1.37 x 1010 = 9.7 x 109

nt = (0.58k/0)Ra'2 (Eq. 7.25 Holman, 1981) (6.5)

0.026 x 0.58 x (9.7 x 109)'2/1.65
4.6 w/m2 - °c

155

QCt = 4.6 x 2.6 x (39-12.4) = 318 u (1,145 kJ/h)
Total furnace loss - Qrt + Qrw + QCt = 8,498 kJ/h

The convective heat loss from the furnace wall is recovered through

the shroud.

3. Duct/bin plenum heat loss:
The heat loss from the duct connecting the furnace to the bin plenum is
computed by the product of the temperature difference between the
diluted air temperature at the furnace exit and in the dryer plenum,

the air mass flow rate (kg/h), and the specific heat of air. Thus:

oduct MaCaAT (6.6)

18,781 x (99 - 77)

413,182 kJ/h

where Ma is the drying air mass flow rate (kg/h); Ca is the specific
heat of air (kJ/kg°C); and.8Tis the difference between the drying air
temperature at the furnace (after the flue gases are diluted with the
ambient air) and in the drying bin plenum (°C). The total system heat
loss is equal to 421,680 kJ/h, the sum of the total loss from the
furnace (8,498 kJ/h) and the combined loss from the duct and the bin
plenum (413,182 kJ/h).

The total energy required is 1,635 MJ/h, being the sum of the
energy required to raise the air from the ambient to the drying tempera-

ture (1,213 MJ/h) and the heat losses (421.68 MJ/h). The furnace/dryer
system efficiency is therefore 74.2% (100 x 1213/1635) at these

156

conditions. Using an average heating value of 20,965 kJ/kg of wood,
the required fuel feed rate is 78 kg/h of wood chips dry matter, which
is equivalent to 92 kg/h of wet wood chips at 14.6% moisture (wet basis).
The experimental and simulated fuel feed rates (at 14.6% moisture)

are 85 and 80 kg/h, respectively (see Table 6.1).

In Table 6.2 the simulated and experimental heat loss values
are compared. The simulated heat loss is computed using the procedure
outlined in Chapter 4. The experimental and the simulated heat losses
are similar. The experimental loss is greater than the simulated heat
loss due to the following reasons:

1. The experimental flow rates of the air mass are higher than
the simulated flow rates. The experimental volumetric airflow rates
are determined from the characteristic fan curve using the experimental
pressure data taken in the drying bin plenum. The volumetric flow rate
is converted to mass flow rate by dividing the former by the specific
volume of air as a function of the dry bulb temperature. However,
the temperature in the bin plenum may not have been a true representa-
tive of the average temperature in the bin (Silva (1980). Empirical
equations were developed to compute the total airflow through the
furnace and into the dryer bin plenum as a function of the bin plenum
pressure and the fuel feed rate. The equations are based on linear
regression analysis of the experimental pressure and air flow data.

The relation between Pf, the furnace fan pressure, and Pb’ the pres-

sure in the bin plenum in kPa is:

Pf = 0.3076Pb + 0.3058 (6 7)

157

The furnace volumetric air flow (m3/min) was determined from the
characteristic fan curve using the pressure Pf obtained from equa—
tion (6.7). The mass flow rate (kg/h) is determined by dividing the
volumetric flow rate by the specific volume at ambient temperature.
The equation for determinine the velocity of air blown around the

furnace by the dryer fan is:

= 146.28 - 13.08P (6.8)

Ff b

where Ff is the air flow velocity around the furnace (m/min). The

mass flow rate of the air around the furnace is determined by multi-
plying the velocity by the cross-sectional area of the annular space
between the shroud and the furnace, and dividing the product (m3/min) by
the specific volume of air at ambient temperature.

The total gas flow out of the furnace is the sum of the mass
flow rate of the air around the furnace and through the eductor, and
the flue gas from the combustion process, The mass flow rate of the
flue gas is the sum of the combustion air mass and the fuel rate, less
the ash fraction in the fuel.

The predicted drying mass air flow is equal to the total gas
flow out of the furnace calculated by the above procedure. The total
air mass flow computed by the above method is slightly less than the
experimental values (as shown in Table 6.2). This is due to leakage
of air into the dryer fan through the duct section between the furnace
and the dryer fan. The empirical equations do not account for the

air leakage of the system.

158

2. The experimental temperature difference between the diluted
air after the furnace and in the bin plenum may be in error due to

the location of the thermocouples.

If the experimental mass airflow rate and heat losses for Test
No. 1 are used, the expected fuel feed rate is 92 kg/h. The experi-
mental fuel feed rate is 85 kg/h, compared to the simulated fuel feed
rate of 80 kg/h. ‘ Thus, for Test No. 1, the experimental method of
determining the heat loss overestimates the actual heat loss by about
8% and the simulation method underestimates the heat loss by approxi-

mately 6%.

6.2.1 Factors Affecting the Fuel Feed Rate
The fuel feed rate depends upon the following factors:
1. Heat loss from the furnace/dryer system as affected
by the ambient conditions (temperature, wind veloc-
ity, and precipitation);
2. the fuel moisture content;
3. the required drying air temperature and air mass

flow rate.

6.2.1.1 Effect of Ambient Temperature on Fuel Feed Rate

The simulated fuel feed rate, heat loss and the furnace/system
efficiency at varying ambient temperature and at a constant drying air
temperature of 80°C are presented in Table 6-3.

The furnace efficiency is defined as the ratio of the total

 

useful energy out of the furnace to the total thermal energy input into

the furnace. It is calculated by the equations:

159

Table 6.3.--Effect of ambient air conditions on fuel consumption,

furnace and system efficiency, and total heat loss
from CVCF/IBCF drying s stem; drying air tem = 80°C,

 

 

 

 

drying airflow = 12.3 /m2-min.
Ambient conditions
Fuel Heat Furnace System
Temperature Absolute consumption loss efficiency efficiency
°C humidity* kg/h MJ/h %
kg/kg
0 0.0029 115 482 95.9 75.0
2 0.0033 110 462 95.8 74.9
4 0.0038 106 447 95.6 74.9
6 0.0043 102 432 95.4 74.7
8 0.0050 98 417 95.3 74.6
10 0.0057 94 403 95.1 74.5
12 0.0065 90 388 94.9 74.3
14 0.0075 87 377 94.7 74.2 .

 

 

 

 

 

 

*Ambient air absolute humidity at 75% relative humidity.

 

160

01.n = anu - MW) (6.9)
n, = (1 - QL/Ain)100 (6.10)

where Qin is the thermal energy input to the furnace (kJ/h); QL is
the heat loss from the furnace (kJ/h); H is the fuel higher heating
value (kJ/kg dry matter); MN is the fuel moisture content (decimal);
and nf is the efficiency (percent).

The system efficiency is defined as the ratio of the net
energy available for drying divided by the total thermal energy input
to the system:

ns = (1 - 05/01.")100 (6.11)

where DS is the total heat loss from the system (kJ/h). The total heat
loss includes the heat loss from the furnace, from the duct work link-
ing the furnace to the bin, and from the wall of the drying plenum of
the drying bin.

The heat loss from the furnace/dryer system increases linearly
with a decrease of ambient temperature. A 1°C decrease in ambient
temperature results in a 7.5 MJ/h increase in the heat loss. The fuel
feed rate required t‘omaintain the same drying air temperature increases
by 2 kg/h (equivalent to 33.5 MJ/h) with the decrease of 1°C in ambient
air temperature. Therefore, the increase in energy input (in the form
of fuel) is higher than the increase in the heat loss.

The increase in the fuel consumption is mainly due to the

energy required in heating the drying air from a lower ambient

161

temperature. Additional energy is utilized in preheating the combus-
tion air and the fuel. The energy required to counterbalance the
heat loss is approximately 22% of the total fuel rate consumption
increase.

According to the simulated results in Table 6.3, the heat loss
varies from 377 MJ/h to 482 MJ/h when the ambient temperature decreases
from 14 to 0°C. According to Silva (1980), the heat loss from the IBCF
(when propane gas was employed) was 351 MJ/h at an ambient and a drying
temperature of 10 and 71°C, respectively: the system efficiency at
these conditions is 71%. The simulated heat loss from the entire CVCF/
IBCF system is 403 MJ/h at an ambient and a drying temperature of 10
and 80°C, respectively, with a system efficiency of 75% (Table 6.3).

The heat loss in a biomass fueled system is slightly higher
than in a propane fueled system due to the extra duct linking the
furnace to the fan. There is no data available to compare the heat
lossfrom a propane fueled system with the biomass fueled system for
the full range of ambient temperatures (0 to 14°C) included in Table
6.2.

According to Silva (1980), the heat loss from the propane
fueled IBCF is equivalent to approximately 5.6 litres/h of liquid
propane (at the ambient and drying temperatures of 10 and 71°C,
respectively). The heat losses in Table 6.3 are equivalent to 22-25 kg
kg of wood per hour and 30-32 kg of corncobs per hour at an ambient
temperature of 14 and 0°C, respectively, when the fuel moisture con-

tent is 20%. In monetary terms the heat losses are equivalent to

162

approximately $1.00/h and $1.26/h when using biomass and liquid pro-
pane fuels, respectively.

The furnace and the system efficiencies are nearly constant
for the 0 - 14°C temperature range presented in Table 6.3. The
furnace efficiency is 95.9% at 0°C and 95.7% at 14°C. The system
efficiency varies from 75% at 0°C to 74.2% at 14°C. These small
changes are due to the fact that the drying temperature is constant,
and therefore, the heat loss increase is compensated by an increase
in fuel feed rate. The furnace efficiency is approximately 95% for
the conditions in Table 6.3. Thus, the furnace heat loss is only 5%

of the total heat loss.

6.2.2 Effect of Fuel Moisture on Fuel Consumption

The fuel moisture greatly influences the average performance
of the CVCF and the required fuel consumption rate at a specific heat
output. The moisture in the fuel reduces the net heating value of
the fuel and the flame temperature, and therefore results in a high
fuel consumption. Incomplete conbustion is: an additional undesirable
effect of low flame temperature.

Table 6.4 lists the simulated heat loss, flame temperature,
fuel consumption rate, and furnace/system efficiencies, when employing
wood chips at different fuel moisture contents and different drying
air temperatures. The simulations are based on an average airflow
rate of 13 m3/min-m2, and an ambient temperature of 4°C. The results

are also plotted in Figure 6.1.

163

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

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165
The three curves in Figure 6.1 fit an exponential curve:

Fr = aebM (6.12)

where Fr is the fuel feed rate (kg/h) and M is the fuel moisture
(percent). The regression coefficient b is 0.018 for the three curves.
The coefficient a depends upon the higher heating value of the fuel

and the required total heat output (as a function of the drying air-
flow rate and temperature as well as the ambient temperature). The
values of the coefficient a for the conditions in Table 6.4 for wood
chips are 60.18, 74.69, and 91.14; the values for corncobs are 67.45,
83.50, and 102.45 for average drying temperatures of 67, 80, and 94°C,
respectively. The coefficients a for corncobs are higher than for wood
chips (at the same drying air temperature) because corncobs have a
lower heating value than wood chips. The coefficient of determination
for all the curves is 0.989.

It can be concluded that the fuel feed rate required for a
given drying condition increases exponentially with the fuel moisture
content. The increase in fuel feed rate is attributed to the reduced
dry matter of the fuel and the heat required to evaporate moisture
from the fuel. The increase in the fuel dry matter feed rate at a
constant drying air temperature is relatively small compared to the
increase in the wet fuel feed rate. At a constant drying air tempera-
ture of 67°C, the wood chips dry matter consumption increases from 67
kg/h when the moisure content is 10% to 76 kg/h when it is 50%.

Table 6.4 also contains information on the furnace performance

characteristics at varying fuel moisture contents at a constant fuel

166

(wood chips) feed rate of 112 kg/h. The results indicate that for
every 10 percent increase in the fuel moisture content, the furnace
temperature decreases by an average of 62°C, and the drying air tem-
perature by 10.6°C. This is due to the decrease in the fuel heating
value since higher moisture fuel contains less dry matter and, there-
fore, less net energy than the same quantity of the lower moisture
fuel.

At a constant feed rate (and constant ambient temperature), the
heat losses decrease inversely with an increase in the fuel moisture
content. A 10% increase in fuel moisture results in 77 MJ/h decrease
in the heat losses. This can be attributed to the lower furnace tem-
perature.

The furnace efficiency remains constant, but the system effi-
ciency at constant fuel consumption increases with an increase in the
moisture content due to a reduction of the duct heat losses. The
duct heat losses decrease since the drying (and therefore;the-duct)
temperature decreases. The increase in the system efficiency is 6%

when the fuel moisture content increases from 10 to 50%.

6.2.3 Effect of Excess Air on Furnace Performance

Excess air is the quantity of the combustion air in excess of
the stoichiometric air required for complete combustion. Some excess
air is necessary in biomass furnaces to ensure complete fuel combustion
(Babcock and Nilcox, 1978; Buchele et al., 1981). If combustion air
is limited to the stoichiometric quantity, incomplete combustion may

result due to a lack of thorough mixing of the fuel and the air. High

167

moisture fuels require high levels of excess air for drying the fuels
prior to combustion (Tillman et al., 1981).

Too much excess air slows down the combustion reaction by
excessive cooling of the furnace combustion chamber. More importantly,
the excess air increases the pressure in and the mass flow rate from
the combustion chamber. This results in backfiring (an explosion of
prematurely ignited fuel) in the feed system and in an increase of
the furnace gas velocities. Unburned entrained materials are carried
out of the furnace and into the drying bin under such conditions.

In the MSU concentric vortex-cell furnace, the combustion and
the eductor air are supplied by the same fan. Increasing the excess
air reduces the eductor airflow. Consequently, the eductor develops
less vacuum in the chamber causing backfiring and smoke-escape through
the feed hopper.

Table 6.5 contains the simulated temperatures in the combustion
chamber and of the drying air along with the furnace and the system
efficiencies at different excess combustion air percentages. The data
are based on an ambient temperature of 4°C and an air flow rate of
12.5 m3/m1n-m2 (17,600 kg/h).

The data in Table 6.5 are plotted in Figures 6.2 and 6.3.
Figure 6.2 is a plot of the furnace combustion chamber and exhaust
flue gas temperatures versus excess air (percent). The furnace and
the system efficiencies as well as the drying air temperature are
plotted against percent excess air in Figure 6.3. The furnace com-

bustion chamber temperature decreases exponentially from 1,518 to

168

Table 6.5.--The theoretical effect of excess combustion air (percent)
on the CVCF furnace performance.

 

 

 

 

 

 

 

 

 

 

 

Furnace and drying Furnace and system
Excess Heat air temperatures efficiency
air loss °C %
% MJ/h
Combustion Exit Drying Furnace System
0 591 1518 579 82.6 92.9 72.3
10 510 1457 590 82.9 93.6 72.9
20 499 1400 600 83.2 94.3 73.4
40 481 1300 615 83.5 95.4 74.4
‘50 472 1248 621 83.5 95.8 74.9
60 464 1203 626 83.4 96.2 75.3
80 446 1117 631 82.9 97.0 76.2
100 427 1038 632 82.0 97.7 77.3
140 379 896 622 79.0 99.2 79.8
180 345 763 594 74.5 99.4 81.6
NOTE: Ambient temperature 4°C, relative humidity 75%, drying
airflow 12.5 m3/min-m2 (292 kg/min)

 

169

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171

763°C when the excess air is increased from zero to 180 percent.
The furnace temperature versus percent excess air curve in Figure 6.2

fits the exponential equation:

Ff = aebZ (6.13)

where Tf is the furnace (combustion chamber) temperature (°C) and

Z is the excess air (percent). The constants a and b are regression
coefficients which depend on the fuel feed rate, fuel moisture, fuel
heating value, and the ambient temperature. The values of the coeffi-
cients a and b are 1,512.7 and 0.004 for the conditions presented in
Table 6.5. The coefficient of determination is 0.9999.

The decrease in the furnace temperature (at constant fuel feed
rate and ambient temperature) is due to the fact that the energy in
the fuel is used to heat a larger air mass.

The exit flue gas temperature increases from 579 to 632°C when
the excess air increases from 0 to 100%, and then decreases to 594°C
at 180% excess air. The set of exit flue gas temperatures and the
percent excess air data (also plotted in Figure 6.2) fits a cubic

polynomial of the form:

2

Te = a + bZ + Cl + dZ3 (6.14)

The regression coefficients a, b, c, and d depend upon the fuel feed
rate, moisture and heating value, and the ambient temperature. For

the conditions presented in Table 6.5, least squares regression yields

172

the following values of the coefficients (coefficient of determination
equals 0.9998): a, 578.86; b, 1.17; c, -6.78 x 10's; and d, 4.15 x
10'5.

The variation of the drying temperature with an increase in
excess air is similar to that of the exit flue gas temperature as
shown in Figure 6.3. It can be expressed in the form of the cubic
equation (6.1.4). The values of the coefficients (if Te is replaced by
Td, the drying air temperature) are: a, 82.56; b, 0.04; c, -4.87E-4;
and d, -8.58E-9. The r2 for this fit is 1.00. This is evidenced in
Table 6.5: the drying air temperature increases from 82.6 to 83.5°C
when the excess air increases from zero to 50%, and then decreases to
74.5°C at 180% excess air.

The variation of the exit flue gas and the drying temperatures
with an increase in excess air is attributed to the variation in the
eductor air mass flow rate and temperature. The eductor and the com-
busion air are supplied by the same fan via a T junction. Excess
combustion air is increased by opening a butterfly valve built into
the line supplying the combustion air. An increase in the combustion
air mass flow rate results in a decrease in the eductor flow rate
and vice versa.

If the combustion airflow increases (with an increase in the
excess air), the temperature of the flue gases leaving the combustion
chamber decreases as previously explained. However, the mass flow

rate of the flue gases increases. The temperature of the exit flue

gases (the mixture of the products of combustion and the eductor air)

173

is calculated by the following equation:
T8 = (MfoTf + Mnchn)/(Mfcf + MnCn) (6.15)

where M is mass flow rate (kg/h); C is the specific heat (kJ/kg°C);

T is the temperature (°C); the subscripts e, f, and n denote the '
eductor exit, the furnace, and the eductor-nozzle conditions, respec-
tively. Equation (6.15) shows that if Mf increases, Te and Mn decrease.
This results in an increase or decrease of the exit flue gas tempera-
ture depending on the relative increase of Hf and decrease of Tf and
Mn. When excess air increases from zero to 100%, the increase in Mf
and the decrease in Mn offset the decrease in Tf, resulting in a slight
increase in Te. When excess air increases above 100%, the decrease in
Tf is more dominant and therefore, the exit fue gas temperature (Te)
decreases (see Figure 6.2).

The drying air temperature varies directly with the exit flue
gas temperature if the ambient temperature remains constant. Therefore,
the drying air temperature versus excess air curve has a similar shape
as the exit flue gas temperature curve. The drying air temperature
decreases faster than the exit flue gas temperature (excess air greater
than 100%) due to heat losses between the furnace and the drying bin.

The plots in Figure 6.2 and 6.3 indicate that the variations in
the exit flue gas and the drying air temperature are small. A greater
amount of excess air causes a greater mass flow rate of gases leaving
the combustion chamber and less mass-flow through the eductor: the

total mass-flow through and around the furnace remains the same. This

174

is due to the fact that the combustion air and the eductor air are
suppllied by the same fan, as mentioned earlier. Excess combustion air
can only be increased by reducing the eductor flow rate. Since the
total airflow through the furnace remains constant, at a constant fuel
feed rate (constant energy input) and drying airflow rate, the drying
air and the furnace exit flue gas temperatures remain nearly constant.

The data in Table 6.5 indicate that the system heat loss
decreases from 591 to 345 MJ/h as the excess air is increased from
zero to 180%. The decrease in heat loss is due to the decrease in the
temperatrue of the furnace combustion chamber and the drying air. Since
the furnace and the system efficiencies vary inversely with the heat
loss, they increase with an increase in escess air. Furnace effi-
ciency increases from 93 to 99% and the system efficiency increases
from 72 to 82% with an increase in excess air from zero to 180%.
Unlike many furnaces (used for generating steam), in which excess
air results in additional heat loss through the stack, all the com-
bustion products are passed through the grain in the MSU CVCF/IBCF
system. Thus, the heat in the excess air is utilized. Moreover,
higher excess air results in lower furnace temperatures, and subse-
quently in reduced radiation and convection heat losses from the
furnace.

A 50% excess air was used in simulating the furnace performance
except when excess air was varied to study its effect on the furnace
performance. The 50% excess air is equivalent to an air/fuel ratio

(dry basis) of 9.5 for wood chips and 8.3 for corncobs.

175

Excess air is difficult to measure without an exact knowledge

of the fuel composition. Excess combustion air is expressed as:
Z = 100(A - S)/S (6.16)

where Z is the excess combustion air (percent), A is the actual
combustion air mass flow rate (kg/h), and S is the stoichiometric
combustion air required (kg/h).

Determination of the stoichiometric air requires an exact
knowledge of the fuel composition (e.g., the fuel ultimate analysis).
It is not possible to describe the fuels used on a farm at any given
time in terms of the exact ultimate analysis without actual measure-
ment. Corncobs may contain randomly varying amounts of husks, corn
kernels, bees wings, and other foreign materials. Thus, the ultimate
analysis of cornbobs as found in the literature represents an average.
The same argument holds for wood chips obtained from a sawmill; the
sample may contain several tree species of wood and bark, each of
which has a different chemical composition and heating value. Thus,
the stoichiometric air for such heterogeneous fuels is difficult to
determine.

If the stoichiometric air is unknown or difficult to quantify,
the excess air cannot be measured.

A measurable unit in a biomass furnace is the air/fuel ratio.
The optimum air/fuel ratio (and also the optimum excess air) varies
between samples of the same biomass fuel depending upon the exact
condition of the fuel in terms of moisture content, particle size

distribution, and the foreign material fraction in the fuel. However,

176

as long as sufficient combustion air is supplied to the furnace,
small changes in the air supply do not have a significant effect on
the combustion efficiency and the heat recovery of a biomass furnace

(Sukup et al., 1983). This is evident from Table 6.5.

6.2.3.1 Fuel Higher Heating Value

The higher heating value of the two biomass fuels used in this
study have been measured using a standard bomb-calorimeter method
(ASTM Standard No. 03286-77, Institute of Gas Technology, 1978).
The values are presented in Table 6.6. The average of the heating
values found in the literature is also listed in the table. There is
no significant difference between the two sets of values. The litera-
ture values are more suitable for simulation purposes since they
represent measured values under a wide range of experimental and fuel

conditions.

6.3 In-Bin Counterflow Corn Drying

The results of the corn_drying tests are presented in Table
6.7. Included in the table are the average ambient and drying air
temperatures (°C), the initial, intermediate, and final grain moisture
contents, the electricity and the biomass fuel usage (electricity
kN/h, fuel kg/h), the specific energy consumption (SECO, kJ/kg of
water removed as defined in section 6.3.2), the corn drying capacity
(wet ton/h), and the drying energy costs ($/wet ton).

The average ambient temperature varied from 2 to 13.5°C in the

fall of 1983, and from -0.9 to 13.3°C in the fall of 1982. The average

177

Table 6.6.--Experimentally measured higher heating value (Md/kg) of
wood chipsand corncobs at different moisture contents
(percent wet basis).

 

 

 

 

Fuel Moisture Heating values, MJ/kg
content % as is dry basis
Hood chipsa 13.3 17.46 20.15
14.4b 17.57 20.51
41.3b 11.35 12435
20.13 .(Average)
20.97 (LA)°
Corncobs 13.05 16.56 19.05
16.43b 15.00 11425
18.50 (Average)
18.65 (LA)c

 

 

 

 

aMood chips consisted of a mixture of soft maple (50%), aspen
(40%) and 10% of other species such as hard maple and unidentified
bark.

bSamples artificially wetted.

cLA = Average of values found in literature (see Table 2.2).

 

178

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179

ambient relative humidity varied from 70% to 100% during the two test
periods.

In Test No. 1 (Table 6.7) corn was dried from an initial
moisture content of 26.2% to an intermediate moisture content of 18.8%
in the high-temperature IBCF system and dumped hot(about 60°C) in the
dryeration bin. The corn was held for approximately 6 hours after
which it was cooled to the 16% final moisture content using ambient
air at a low airflow rate (approximately 0.5 m3/min-m2 per m3 of corn).
The corn was similarly dried in the Tests Nos. 2, 3, and 5. During
the Test Nos. 4, 6, and 7 the corn was dried directly to a final
moisture contents of 14.2%, 13.1%, and 14.6%, respectively, in the
high-temperature system.

The electric power consumption was 10.7 kWh/h during the
heating-drying phase. The difference between the 10.7 kWh/h and

the rates in Table 6.7 are due to the electrical energy required to

operate the cooling fan in the dyeration bin.

6.3.1 Biomass Fuel Feed Rate

The average fuel feed rates are included in Table 6.7. The
fuel feed rate depends upon the total thermal energy required for the
drying process as discussed in section 6.2 (see also Table 6.1 for the

same information).

6.3.2 Experimental Specific Energy Consumption
The specific energy consumption (SECO, kg/kg of water removed)

is presented in Table 6.7 under two categories, namely the net SECD

180

(for drying only) and the gross $600 (which includes the heat loss
from the system). The net SECD is calculated by multiplying the tem-
perature difference between the ambient and the drying air (°C) by
the specific heat of air (kJ/kg) and the mass flow rate of the drying
air (kg/h), adding the product to the electricity used (kJ/h), and
dividing the sum by the water removed from the grain (kg/h). The
gross SECO is calculated by multiplying the fuel dry matter consump-
tion by the higher heating value and adding the electrical energy to
the product; the sum is divided by the amount of water removed.

As can be seen in Table 6.7, the net and gross SECO depend
upon the combined effects of the following factors:

1. The amount of water removed per ton of wet corn.

2. The moisture content percentage points removed after

drying in the high-temperature phase.

3. The ambient temperature.

4. The drying temperature.

5. The airflow rate of the drying air.

The effects of the above factors on the drying capacity and
the SECO are discussed in detail in section 6.5. It is obvious that
when drying is completed in the high-temperature phase (i.e., corn
dried to 15.5% or less), the process results in high SECO values. Thus,
Tests 4, 5, and 7 resulted in high SECO values compared to Test No. 1.

The ambient and the drying air temperatures, as well as the
drying airflow rates, have a direct effect on the SECD values. This
effect is not obvious from Table 6.7 since some of the important para-

meters varied simultaneously.

181

6.3.3 Drying Capacity

The average drying capacity at an average drying temperature
of 84°C for the seven tests shown in Table 6.7 is 2.2 wet tons per
hour in drying from 25% to 15%. The drying capacities for Tests 6
and 7 are much lower than for the other tests (1.5 and 1.8 compared
to 2.7 and 2.5 wet tons per hour for Tests 1 and 2, respectively).
This is due to the fact that the corn was dried to a very low mois-
ture content in the high-temperature phase (14.6% and 14.7% compared
to 18.8% and 17.0% in Tests 1 and 2). If Tests 6 and 7 are excluded,
the average capacity for the other tests is 2.4 wet tons per hour.
This compares with a capacity of the IBCF fueled with propane of 2.6
wet tons per hour in drying from 26% to 17% at a drying air temperature
of 71°C (Silva, 1980). Thus, there is a small loss in drying capacity
when the IBCF is operated with biomass fuel. The slight decrease in
capacity is due to the air flow reduction in heating the air before the

fan entrance.

6.3.4 Standardized Dryer Performance

The standardized dryer performance data are presented in Table
6.8. The data include the value of the energy consumption (fuel hg/ha,
electricity kWh/ha), capacity (wet ton/h), SECO (kJ/kg of water
removed), and the drying costs ($/wet ton). The standardized per-
formance is based on drying corn from 26% to 18% moisture in the
high-temperature phase and from 18% to 15.5% in the dryeration/cooling
bin. The energy consumption per hectare is based upon a corn yield of

6.28 wet tons per hectare.

182

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183

The following is the standardization procedure:
1. The corn drying constant k for each test is derived
from the drying equation:

M - M
MR = fi = exp(-kt) (6.17)

Rearranging equation (6.17):
k = -tn(MR)/t (6.18)

2. The theoretical time required to dry corn from

26 to 18% moisture is calculated as follows:

tS = £n[(0.IB-Me)/(0.26-Me)]/k (6.19)

3. The standardized capacity (wet ton/h) is:

CPs = CPe/tS (6.20)

where CPS and CPe are the standardized and the experimental capacity
(wet tons per hour). The other standardized data are based on the

standardized capacity (CPS) and time (ts).

6.3.4.1 Standardized Drying Capacity

The standardized capacity is 2.2, 2.5, and 2.7 wet tons per
hour for a drying air temperature of 78, 82, 89, and 93°C, respectively.
The average standardized capacity is 2.5 wet tons per hour at an inlet

air temperature of 84°C.

184

6.3.4.2 Standardized Specific Energy Consumption

The average standardized gross SECO for the CVCF/IBCF system
is 6,099 kJ/kg of water, compared to 4,548 kJ/kg if propane is used
(Silva, 1980). The difference is due to the extra heat loss through
the additional duct work linking the furnace to the dryer (see section

6.2).

6.4 Corn Drying Simulation

The in-bin counterflow grain drying model discussed in section
4.2 was used to simulate the drying of corn in the CVCF-IBCF system.
The simulated and experimental results are compared in Tables 6.9 and

6.10.

6.4.1 Cycle Times

The in-bin counterflow dryer (IBCF) is a semi-continuous
counterflow dryer. Grain is dried in a stationary mode as in an in-bin
batch drying system. When a thin layer (7.5-10 cm deep) of corn at
the bottom of the bed is dried to a predetermined moisture content, it
is transferred to a dryeration bin (see section 4.2). The transfer
process takes approximately 25 minutes during which the drying
proceeds. The time between two consecutive starts of the transfer
process is defined as the cycle time. It is the time required to dry
the 7.5 - 10 cm of grain that is transferred to the dryeration bin.

The experimental and simulated cycle times, moisture content
at transfer, capacity (wet tons per hour), airflow rate and specific
energy consumption rate are tabulated in Table 6.9 for Test No. 1.

(See the tables in Appendix B for an explanation of other tests.)

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186

Table 6.10.--Simulated operating conditions during Test 1.

 

 

 

 

 

Bottom layer
Exhaust air moisture content Average Remaining
Cycle % moisture bed
No. Temperature RH in the bin depth
°C % Before After: % m
cycle cycle
1 26 100 26.2 18.7 26.2 1.85
2 38 100 19.9 18.5 26.8 1.69
3 43 100 20.7 18.8 27.0 1.53
4 44 100 20.9 18.6 26.5 1.37
5 45 100 20.6 18.8 26.0 1.21
6 47 92 20.5 18.8 25.4 1.05
7 48 65 20.3 18.6 24.7 0.09
8 54 42 20.2 18.6 23.8 0.74
9 59 27 20.1 18.5 23.0 0.58
10 64 18 20.0 18.8 22.5 0.42
11 70 10 20.0 18.8 21.5 0.26

 

 

 

 

 

 

 

 

187

The cycle times at different cycles are plotted in Figure 6.4. The
first cycle requires the most time. The first cycle time is the time
duration from the moment the drying fan is started until the start of
the transfer of the first dried layer of corn to the dryeration bin.
It requires more time to dry the first layer because of the following
reasons:

1. The grain is initially cold, and it requires time

(and energy) to heat the grain (and the drying struc-
ture) before evaporation of moisture from the grain
is initiated.

2. All the grain in the bin is nearly at the same high

initial moisture content and low temperature. More
time is required to dry the initial bottom layer to a
specified moisture content than is the case for sub-
sequent layers.

3. The airflow increases as the bed depth decreases.

The length of the time interval between the start of the dry-
ing process and the first transfer depends upon the initial grain
temperature and moisture content, the grain depth, the drying tem-
perature, and the final (transfer) moisture content.

The cycle time of the second transfer is shorter than of the
others. During the drying of the first cycle, the grain above the
layer transferred is dried to a moisture close to the desired final
moisture content. Thus, it takes a short time to dry the grain.

The simulated average moisture contents of the transferred corn

before and after drying, the average moisture of the grain left in the

188

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189

bin, and the exhaust drying air temperature and relative humidity

are presented in Table 6.10. It can be seen that the initial mois-
ture content for cycle 1 is 26.2%; for cycle 2, 19.9%; and for cycle 3
to 9 the initial moisture is about 20%.

As the drying continues, the following changes occur:

1. The average grain temperature is increased.

2. The average moisture content of the grain remaining

in the drying bin first has increased due to rewetting
of the grain at the top, and then is decreased as the
drying front moves upward (see Table 6.10).

3. The airflow rate is increased since the resistance to

the airflow is decreased as the grain depth decreases.

4. The initial moisture content for each cycle (being

the moisture content at the bottom layer after removal
of dry grain in the previous cycle) is decreased
gradually.

5. The time between consecutive cycles decreases and

therefore, the drying rate (capacity) is increased
due to the combined effect of factors 1-4 above.

The experimental results in Tables 6.9 and 6.10 show some
deviations from the expected trends, such as the experimental cycle
time for cycle No. 10, which increased to 1.05 hrs compared to the
previous and the following cycle times of 0.80 and 0.93 hrs,
respectively. Also, the experimental cycle time increased from 0.77
hrs for the fifth cycle to 0.95 hrs in the 7th cycle. These irregular-

ities are due to the falling drying air temperature, the difference in

190

the original moisture content of grain in different layers, or a

combination of the two.
6.4.2 Drying Capacity
The drying capacity (wet ton/h) is defined as

Capacity = Gilt (6.21)
° 1

'l

IIMZ

where N is the total number of cycles dried after time t (hours) and
Gi is the weight of wet grain dried in cycle i (ton). The average
capacity can also be defined as the product of the number of cycles
and the average amount of grain dried per cycle, divided by the total
drying time.

Both the simulated and the experimental capacities increase
with time (and therefore with the number of cycles) as the cycle times
decrease. The reasons for the reduced cycle time are discussed in
section 6.4.1

There is close agreement between the simulated and the experi-

mental capacities (3.0 and 2.7 wet ton/h, respectively).

6.4.3 Specific Energy Consumption

The simulated specific energy consumption (SECO) presented
in Table 6.9 is the net energy required to remove one kilogram of
water from the wet grain. The net SECO does not include heat losses

from the system and is computed as follows:

191

SECO( =

N) (GaiCATAti + Ei)/Wi (6.22)

IIMZ

i 1

where Gai is the airflow during the drying of cycle i (kg/h); C is

the specific heat of air (kJ/kg°C); AT is the temperature difference
between the ambient and the drying air; Ati is the cycle time (h);

E1 is the electrical energy used in drying cycle i (kJ); and W1 is the
amount of water removed in cycle i (kg).

The total amount of water removed after a given cycle is

calculated as follows:

”i ‘ (moi ' mfi)GDi + (mBo ‘ mBi)GBi (5°23)

where moi is the initial moisture content of the grain at the transfer
layer dried in cycle i (decimal, dry basis); mfi is the final moisture
content of the grain transferred in cycle i; mBO is the initial average
moisture content in the bin above the layer removed in cycle i;

mBi is the average moisture content of the grain remaining in the bin
after cycle i has been transferred; GDi is the dry matter of the corn
transferred in cycle i; and GBi is the dry matter of the corn remaining
in the bin after cycle i.

The specific energy consumption at the end of each transfer as
defined by equations (6.22) and (6.23) has not been determined
experimentally. It requires the grain moisture content of the bottom
layer before and after the layer has been dried as well as the average

moisture of the grain remaining in the bin. These values cannot be

obtained in a commercial operation.

192

The variation of the simulated specific energy consumption
as drying progresses is illustrated in Table 6.9. The SECO is low
during the first cycle, reaches a maximum during the second cycle,
and decreases until the 8th cycle is reached.

After drying is started, 1.43 hours elapses (1.8 hours expe-
rimental) before the first layer of dried grain (26-18.7%) is dis-
charged from the dryer. During this time the grain immediately above
the bottom layer is dried to 19.9%. Thus, the initial moisture for
the second cycle is only 19.9%, compared to 26.2% for the first cycle.
Therefore, considerably more water is removed in the first cycle than
in the second (224 kg in the first cycle compared to 86 kg for the
second cycle). The shorter drying time for cycle No. 2 results in a
higher specific energy consumption because there is little water
removed in drying the corn from 19.9 to 18.5%. Due to the shorter
drying time for the second cycle, the grain above the layer transferred
is 20.7% (compared to 19.9% for the second cycle).

The third cycle requires 0.76 hours to dry. The fourth cycle
time takes longer than the third cycle for the same reasons. However,
after the fourth cycle, the cycle time progressively decreases because
the grain is sufficiently heated namaintain a steady drying rate.

After the third cycle, the amount of water removed per cycle
remains nearly constant (143 kg/cycle). However, the cycle time
progressively decreases since the average temperature of the corn is
also increasing as the drying continues. The decrease in the energy

comsumption is shown in Figure 6.5.

193

 

 

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2
2

 

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Figure 6.5. Simulated Specific Energy Consumption as a Function of Number

194

As drying progresses, the grain depth also decreases. The
decrease in depth results in an increase in airflow rate. After the
6th cycle the exhaust air leaving the drying bin is no longer saturated
(see Table 6.10). The exhaust relative humidity decreases and the
temperature increases. This is an indication of the start of ineffie
cient use of energy. The specific energy consumption reaches a minimum
after the 8th cycle and thereafter increases. The increase is a result
of the decreasing amount of water removed per cycle. The depth at
which the minimum specific energy consumption is reached depends upon
the drying air temperature, the initial grain depth and moisture, and

the degree of drying.

6.4.4 Airflow Rate

The simulated and the experimental airflow rates at the begin-
ning of each cycle are tabulated in Table 6.9 and plotted in Figure 6.6.
The experimental airflow rates were determined by measuring the static
pressure (in the bin plenum) and the bed depth. The characteristic
fan curve was also used to determine the airflow rate using the static
pressure readings.

The simulated airflows were predicted employing the procedure
discussed in section 4.2. There is a close agreement between the

experimental and simulated airflow rates.

6.4.5 Comparison of Experimental and Simulated Grain Moisture Contents
Experimental and simulated moisture contents of each cycle are

presented in Table 6.9.

195

 

 

 

 

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Figure 6.6. Simulated Airflow Rates at the Beginni

196

The desired final moisture content is one of the inputs to
the grain drying model. Whenever the average grain moisture content
within the bottom layer (7.5-10 cm above the false floor) reaches the
desired moisture, the layer is removed from the drying bin instantane-
ously and dumped in the cooling bin. Therefore, the simulated mois-
ture content is always equal to or slightly less than the desired value.

The difference between the average experimental and simulated
moisture contents is small (e.g., 18.8% experimental compared with

18.7% simulated for Test No. 1, Table 6.9).

6.4.6 Comparison of Experimental and Simulated Drying Data

Table 6.11 contains the experimental and simulated corn drying
results for seven tests under different drying and ambient conditions.
The average drying capacities and net specific energy consumption

results are compared.

6.4.6.1 Capacity

The experimental capacity was determined by dividing the total
amount of corn dried by the total drying time. The simulated capacity
was determined by the procedure outlined in section 6.4.2. The data
indicate close agreement between the experimental and simulated
results.

The difference between the experimental and the simulated
capacity is due to the following factors:

1. The sweep auger does not remove equal amounts of

grain each time when grain is transferred to the

197

 

 

 

 

 

 

 

 

 

 

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198

dryeration/cooling bin due to uneven grain flow
characteristics such as density, broken kernels,
foreign materials (BCFM), etc.

2. The simulation model uses average ambient and
drying temperatures, and an average initial grain
moisture content; these parameters vary somewhat

during a test.

6.4.6.2 Simulated Versus Experimental Specific Energy Consumption

The experimental and simulated net specific energy consumption
values for the heating and cooling phases are also presented in Table
6.10. The experimental SECO data are based upon the average airflow
rate, the average moisture content, and the average drying capacity.
The simulated SECO values are based on the integration of the SECO
values for every cycle as discussed in section 6.4.3.

There are differences between the experimental and the spe-
cific energy consumption values due to the methods used to compute
them. The integration method (used to calculate the simulated SE00)
is more accurate but requires a knowledge of the average moisture
content of the grain in the bin, as well as the moisture of the grain

at the bottom of the bin before and after each transfer.

6.5 Predicted Drying Performance Parameters
The furnace and the drying models are used to predict the
furnace/dryer performance under varying ambient conditions and dryer

operating modes. The IBCF dryer is operated in four basic modes:

Mode 1:

Mode 2.

Mode 3.

Mode 4.

199

The bin is filled to a depth of 1.8 m. Drying is started
and continues until the bin is empty. A total of 30 wet
tons are dried in 8 to 12 hours depending on the drying
temperature, the initial moisture content, and on whether
drying is completed in the high-temperature phase or in
the cooling bin.

The bin is initially filled to a depth of 1.8 m. Drying
is started and the bin is refilled with 8 tons of wet
grain every three hours. Drying continues until grain
loading is stopped after 18 hours. A total of seven batches
are added during the drying period. Approximately 57

wet tons are dried (21 cycles).

As in Mode 2 except that 16 wet tons per batch are added.
Drying continues in this mode until the bin is filled up
to 3.7 m after 11 hours. Four batches (a total of 64 wet
tons) are added to the bin. A total of 12 cycles (32 wet
tons) are dried.

As in Mode 2, but 4 wet tons are added per batch. The
drying rate is greater than the refill rate. Drying con-
tinues until the bin is empty after 16 hours. A total

of five batches (20 wet tons) are added to the initial

30 wet ton. Fifty wet tons (17 cycles) are dried during
the 16-hour drying period.

In addition to the four modes discussed above, the IBCF dryer

is used to partially dry grain to an intermediate moisture content and

200

finally dry the grain in the dryeration bin (Method 0). Alternatively,
the grain can be dried directly to the final moisture in the IBCF
dryer (Method ND).

Table 6.12 presents the predicted IBCF system performance
parameters (simulated) in different drying modes and at different
drying temperatures. The following parameters are studied: capacity,

specific energy consumption, fuel feed rate, and drying costs.

6.5.1 Effect of Refill on Dryer Performance

The IBCF performance is better when operating in Modes 2 and 3,
than in Modes 1 and 4. The capacity is 3.2 and 3.1 wet ton/h in
Modes 2 and 3, respectively, compared with 3.0 and 2.9 wet ton/h in
Modes 1 and 4. The net specific energy consumption is 4656 and 4792
kJ/kg of water removed in Modes 2 and 3 compared with 5010 and 5126
kJ/kg in Modes 1 and 4. The operating (energy) costs are also lower
in Modes 2 and 3 than in Modes 1 and 4 ($1.81, $1.86, $1.90, and $1.91,
in Modes 2, 3, 1 and 4, respectively).

The performance is superior in Modes 2 and 3 than in Modes 1

or 4 for the following reasons:

1. Replenishing the bin with wet grain ensures complete
saturation of the exhaust air. Thus, the air water-
carrying capacity is fully utilized, resulting in
better energy utilization.

2. The wet grain is preheated by the exhaust air and

therefore, dries faster upon reaching the bin bottom.

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202

3. There is no increase in the airflow rate since the
depth either remains constant (Mode 2) or increases
with time (Mode 3); the airflow rate increases in

Modes 1 and 4 with a decrease in grain depth.

6.5.2 Effect of Dryeration on System Performance
The data in Table 6.12 indicate that the dryeration process
results in a higher drying capacity, a lower specific energy consump-
tion, and a lower drying energy cost compared to drying completely
in the IBCF dryer. When the drying air temperature is 80°C, the
simulated drying capacity is 3 and 2.6 wet ton/h with and without
dryeration, respectively; the respective SECO values are 5,010 and
5,573 kJ/kg of water, and the energy cost per wet ton $1.90 and $2.16
(wood chip fuel).
The improvement of the system performance with dryeration is
attributable to the following factors:
1. Less moisture is removed by the IBCF dryer resulting
in less drying time of each wet ton of grain in the
counterflow dryer.
2. The last 2.5% moisture remaining in the corn after
the high-temperature drying phase is the portion of
the moisture most difficult to remove. This moisture
is now removed slowly during the aeration process.
3. The aeration process uses a low airflow rate (0.5 -
1 m3/min-m2) and an insignificant amount of electrical
energy to cool the grain and remove the last 2.5% of

moisture.

203

6.5.3 Effect of Drying Air Temperature on System Performance

The drying air temperature has a direct effect on the grain
drying capacity, the specific energy consumption, and the drying
costs. The effect of the drying air temperature is shown in Table
6.12. The drying capacity increases substantially with an increase
in the drying air temperature. The drying capacities are 2.7, 3.0,
and 3.2 wet ton/h when corn is dried from 26 to 18% at the drying air
temperatures of 67, 80, and 94°C, respectively. The higher tempera-
ture air increases the drying rate of the corn and has more water-
carrying capacity since the saturation absolute humidity is higher.

The effect of drying air temperature on the specific energy
consumption depends upon the moisture percentage points removed in
the high-temperature phase. If corn is dried from 26 to 18% the SECD
is 5,010 kJ/kg at 67 and 80°C drying air temperatures, and 5,140 kJ/h
at 94°C. However, if corn is dried from 26 to 15.5% in the high
temperature phase, the SECO is 6,035, 5,573, and 5,972 at air tem-
peratures of 67, 80, and 94°C, respectively. Thus, if drying is com-
pleted in the high-temperature phase (i.e., corn dried down to 16%
moisture content or less), the SECO decreases with increasing drying
air temperature up to a point after which the SECD increases with
the temperature increase. The inflection point depends upon the
ambient temperature, the amount of water removed, and the rate of
increase in drying capacity with the temperature increase.

In the example given above, the drying capacity is 2.2 wet

ton/h at 67°C and 2.6 wet ton/h at 80°C. Thus, the capacity increases

204

by 20%; the energy required'Ulheat the drying air increases by 19%
resulting in an 8% reduction in the specific energy consumption.
When the drying air temperature is raised to 94°C, the capacity
increases by 26% while the energy increases by 40%. Thus, raising
the drying temperature from 67 to 94°C results in a 1% reduction in
the SECO when corn is dried from 26 to 15.5% in the IBCF system.
When corn is dried from 26 to 18% moisture in the high-

temperature phase and finally dried to 15.5% in the cooling phase,
the drying costs are $1.94, $1.90, and $1.96 per wet ton at tempera-
tures of 67, 80, and 94°C, respectively, compared to $2.33, $2.16,
and $2.28 per wet ton if the corn is dried from 26 to 15.5% in the
high-temperature phase. Thus, the drying costs vary inversely with

the drying capacity.

6.5.4 Effect of Initial Moisture on System Performance

Table 6.12 contains the simulated IBCF dryer performance para-
meters when drying corn from 26 to 22% initial moisture contents.
The results indicate that when corn is dried from 22% moisture, the
drying capacity is higher and therefore the drying costs per wet ton
lower. The drying capacity (with dryeration) is 3.8 and 3.0 wet
ton per hour, the drying costs are $1.51 and $1.90 per wet ton for
the 22% and 26% initial moisture content corn, respectively.

The costs per kg of water removed is higher for the 22%
initial moisture corn. The reason for the higher drying capacity
is that less water is removed per ton of wet corn if the initial

moisture is 22% than if it is 26%.

205

Corn with an initial moisture of 22% or less can also be
dried in low-temperature or natural air drying systems (Bakker-

Arkema et al., 1981).

6.5.5 Effect of Ambient Temperature on System Performance

The ambient temperature has a direct effect on the specific
energy consumption and the drying costs. At lower ambient temperature,
the heat loss from the system is greater as well as the energy required
to raise the drying air to a given drying temperature, than at a
higher ambient temperature. The specific energy consumption is 5334
kJ/kg at 0°C ambient temperature, compared to 5010 and 4514 kJ/kg at
4 and 10°C, respectively. The drying costs (wood chip fuel) are $2.01,
$1.90, and $1.77 for drying at 0, 4 and 10°C ambient temperature,
respectively. The drying capacity does not change with ambient tem-

perature if the drying temperature remains constant.

6.6 Economics of the CVCF/IBCF System

The CVCF/IBCF drying system is a technically feasible system.
It would be economically viable if it results in reduced drying costs
compared to a propane fueled IBCF system. A substantial investment
is required to convert the propane fueled system to a biomass system.
The recovery of the additional capital costs must be through savings
in fuel costs.

In this section the fuel costs for operating the CVCF/IBCF
system using wood chip and corncob fuels at different ambient and

drying conditions are presented. The operating costs include the

206

biomass ‘fuel costs, as well as the cost of the electrical energy
required to run the furnace and the dryer fans and the IBCF system
control. The biomass costs are compared with the cost of operating
a propane fueled IBCF system.

A 12-year budgeting analysis is included to show the effect of
fuel costs, labor, repairs and maintenance, taxes, depreciation,
and inflation on the annual break-even costs of operating the CVCF/IBCF
system. The budgeting analysis includes the life cycle costing of
the system. The results are compared with the annual break-even costs

for a propane fueled system.

6.6.1 Operating Costs
The experimental drying costs are presented in Table 6.7. The
costs represent operating costs; they do not include labor and invest-

ment costs. The fuel costs are based upon the folloWing 1983 prices:

 

 

 

Wet wood chips: F.0.B. cost $ 6.00 per ton
Drying and storage 0.00
Freight 0.16 per ton/km

Dry wood chips: F.0.B. cost 15.00 per ton
Drying 3.75 per ton
Storage 2.00 per ton
Freight 0.08 per ton/km

Dry corncobs: F.0.B. cost 25.00 per ton
Drying and storage 0.00

Freight 0.36 per ton/km

207

 

Electricity: 0.07 per kWh
Propane: 0.22 per liter

The cost of biomass fuel depends upon the distance over which
the fuel is hauled. The F.0.B. cost of biomass fuels is made up of
handling expenses incurred in harvesting and/or gathering and storage.
The storage and drying costs are zero for dry corncobs and wet wood
chips. In the case of corncobs, the fuel is gathered during or after
the harvest season and stored in storage bins which are not used for
grain storage during the off-season. The corncobs dry naturally in
storage. The corncob fertilizer value is approximately $25 per ton.
In the case of wet wood chips, the fuel is used directly without
storage or drying.

The cost of dry wood chips includes the drying and the storage
expenses. The fuel is obtained from a saw mill at approximately 50%
moisture and dried to 30% moisture (or less). A special bin is
required for storage and drying by aeration.

The transportation costs depend upon the quantity per load and
the distance. The transportation costs per km (kilometer) for the
corncobs is high because of their low density compared with wood chips.
The bulk density of dry corncobs is approximately 200 kg/m3 compared
to an approximate value of 320 kg/m3 for dry wood chips.

The total costs of the wood chips and corncobs used in this
research are assumed to be $44.70/ton and $35.00/ton, respectively.
High moisture content fuels are more expensive than dry fuels because

the net heating value is much less for the wet fuels. The energy costs

208

(Table 6.7) is $5.82 per wet ton for Test No. 7 compared to $2.37 per
wet ton for Test No. 6. The moisture content of the fuel (wood chips)
for Test No. 7 is 45% compared to 10% for the corncobs used in Test
No. 6. The higher the fuel moisture content, the higher is the fuel
feed rate required to maintain the same drying air temperature.

The average drying cost for the corn in Table 6.7 is $2.74/wet
ton. This compares with $5.75/wet ton for a similarly operated propane
fueled IBCF system (Silva, 1980). Thus, the operating (fuel) costs
for the CVCF/IBCF system are lower than for an equivalent system fueled
by propane (note that the investment and labor costs are excluded).

The standardized energy drying costs are included in Table 6.8.
The cost for drying with 45 and 19.3% moisture wood chips is $3.48 and
$2.45 per wet ton, respectively, compared to $2.25 per wet ton if
14.6% moisture is used; the figure is $3.72 per wet ton if liquid
propane is the fuel.

Table 6.8 shows that dry corncobs are the least expensive
fuel. The cost of drying is $1.58 and $1.68 per wet ton for 10% and
12% moisture corncobs, respectively, due to the low transportation
costs (as discussed in section 6.6.1).

The average standardized energy cost is $2.27 per wet ton for
the conditions in Table 6.8. The costs for similar conditions when
drying with propane fueled system is $3.72/wet ton (Silva, 1980).

Thus, the operating (energy) costs are less for biomass fuel than

for liquid propane.

209

6.6.2 Capital Budgeting Analysis

The operating costs presented in Tables 6.7 and 6.8 do not
include labor, maintenance, investment, interest on borrowed money,
depreciation, and taxes. To analyze these costs, a 380 wet-ton
annual corn drying system was analyzed using a capital investment
model developed by Harsh (1983). The cost estimates are based upon a
simulated energy consumption, drying rate, and operating costs for the
CVCF/IBCF system operating in Mode 1, Method 0 (Table 6.12), at a
drying temperature of 80°C. The relevant input data for the budgeting
analysis are summarized in Table 6.13.

The biomass furnace can be considered an addition to an IBCF
system. The MSU furnace was manufactured at a cost of $11,070; this
includes costs for materials ($8,570), labor ($2,000), and miscel-
laneous expenses ($500). The retail price for the unit is estimated
to be $16,000. The total cost of the CVCF/IBCF system is $43,000
(Kalchik, 1984) compared to $27,000 for the pr0pane fueled system.

It is assumed that the additional $16,000 is borrowed at 12% annual
interest rate and is payable over a five-year period.

The accelerated cost recovery system (ACRS) outlined in the
1983 Famers Tax Guide (IRS, 1983) was applied to calculate deprecia-
tion. The biomass furnace can be classified as a 5—year property
with a recovery period of 12 years when using the ACRS code. A 12-
year, rather than a 10-year, budgeting analysis is presented. A
10—year recovery is not permissible for a 5-year property under the

ACRS method. Also, the expected salvage value of the property is not

210

Table 6.13.--Estimates and assumptions for a 12-year budgeting analysis

(1983) prices) of the CVCF/IBCF drying in Mode 1 (see
Table 6.12).

 

 

 

 

 

 

Fuel
Parameter .
. Wood chips Corncobs

°5tlm°t°° 20% M.C. 20% M.C. Pr°°°°°
Fuel cost
$/ton 1.90 1.72 5.32
$/ton-point 0.18 0.16 0.51
Labor cost
$/ton 1.11 1.16 0.40
$/ton-point 0.11 0.11 0.04
Dryingrate
ton/h 3.00 3.00 3.0
ton-point/h 31.50 31.50 31.5

 

 

 

 

Initial Investment: $27,000 pr0pane, $43,000 biomass
Percent Borrowed: 0.% propane, 35% biomass

Repayment Period: 5 years

Interest Rate: 12%

Tax Bracket: 20

Type of Property: 5 year property

Depreciation Type: Straight line method

Recovery Period: 12 years

Number of Units Dried: 4,000 ton-point per year (381 ton dried from
26-15.5% w.b.)

Losses due to Corrosion, etc.: $120 per year

Repairs & Maintenance: $2,000 propane, $4,500 biomass
Percent Return: 15%

General Inflation Rate: 6%

Fuel Inflation Rate: 10% propane, 2% biomass

Biomass Fuel Moisture: 20% w.b.

Drying Temperature: 80°C

Ambient Temperature: 4°C

Ambient Relative Humidity: 75%

 

211

subtracted from its basic price. The straight line depreciation
method is used to compute the annual depreciation cost as a deductible
item in tax computations under the ACRS system.

The results of the 12-year budgeting analysis are presented
in Table 6.14. The annual operating costs are expressed in $/wet ton-
point. The unit represents the cost of drying one ton (wet) when one
percentage point is removed (e.g., one ton wet corn dried from 26
to 25% MC, wet basis). This unit is preferred over the more common
unit of $/wet ton. The latter unit cannot be readily used to estimate
the drying costs when grain is dried from and to different moisture
contents. The total and per wet ton investment expenditures are
included in the table.

The total annualized break-even drying costs for the biomass
fuels is $3.00 per wet ton-point compared to $2.23 for the propane
system if the annual rate of inflation of the fuel-cost is 2% for all
the fuels. If the pr0pane cost inflation is 10% annually, the total
drying costs for the propane fueled system if $2.26 per wet ton-point.
The latter is considered a more realistic figure. Thus, the propane
fueled system is the least expensive at the current propane price of
$0.225 per litre. The pr0pane cost would have to increase to $0.378
per litre for the total costs of the prOpane system to be equal to
that of the biomass fueled system. This constitutes an increase of
68% in the present propane cost.

The variable annual break-even costs of the biomass system

($0.95 per wet ton-point) are about the same as for the propane fueled

212

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213

system ($1.05 and $0.82 at annual fuel cost inflation of 10% and 2%,
respectively). The total costs for the biomass system are higher due
to the higher investment in the CVCF furnace. The fixed costs are 68%
of the total break-even costs in the case of the biomass system, and
57% in the case of the propane system (with 10% annual fuel cost
inflation).

There is no significant difference between the drying costs for
wood chips and corncobs under the conditions presented in Tables 6.12
and 6.13. As mentioned earlier, the cost of a biomass fuel depends on
the distance from which it is obtained, and the handling costs involved
in drying and storage. Therefore, it should be remembered that the
costs presented in Table 6.14 are affected by the transportation and
handling costs, and thus, the costs would be slightly different if

the transportation and the handling costs changed (see section 6.6.1).

6.7 Corn Contamination in Drying by Direct Biomass Heating

In the MSU concentric vortex-cell biomass furnace, the flue
gases are diluted with ambient air and passed directly through the
drain. The products of incomplete combustion and fly ash are potential
problems with a direct-heated drying system. The primary combustion
products are carbon dioxide, carbon monoxide, ash, soot, aldehydes,
ketones, sulful dioxide, oxides of nitrogen, and polycyclic aromatic
hydrocarbons (PAH). The ash contains inert compounds of heavy metals.
The ash and the soot particles contain a high percentage of the other
combustion compounds and may be condensed and absorbed on the grain

surface.

214

Table 6.15 shows the contamination data of the corn dried with
the MSU CVCF/IBCF system. The following substances were detected in
the corn: naphthalene (Nph), fluorine (F1), anthracene (An), benz
[a] anthracene (BaA), and benzo [a] pyrene.

During the steady-state operation of the MSU biomass furnace,
there is no visible smoke, and only a small amount of fly ash is
observed. The fly ash is filtered out of the drying air by a thin
layer of corn (approximately 3 cm deep) at the bottom of the drying
bin. This layer isa residue which is not transferred to the cooling
bin. It is removed when the bin is cleaned at the end of the drying
season.

The corn dried with wood chips did not contain a measurable
amount of any of the PAH before drying. However, the corn dried with
corncobs did show measurable amounts of fluorine and anthracene before
drying because of its growth environment, possibly close proximity to
a highway (Sailor, 1978).

The PAH concentrations in the biomass dried corn appear to
increase slightly with time of exposure. Corn is subjected to hot
drying air for a maximum of about 8-12 hours in the in-bin drying
system (depending on the grain depth in the bin, the initial moisture
content, and the drying temperature). The PAH concentration was in
the range of 0.06 ppb (BaP) to 8.6 ppb (Nph) after a 6- to 9-hour
drying period. The residual corn at the bottom of the drying bin
showed a considerable higher concentration of anthracene (0.6 ppb,

benz [a] anthracene (1.60 ppb), and benzo [a] pyrene (0.56 ppb) compared

215

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216

to the corn transferred to the cooling bin. The concentration of the
BaA and BaP was less than 0.4 ppb in the corn transferred to the cool-
ing bin.

BaA and BaP have shown some carcinogenic activities in test
animals (Souci, 1968). The range of concentrations of these compounds
is well below that considered to be dangerous to human or animal health
(Anderson et al., 1981). Moreover, since most of the corn dried in
the USA is used for livestock, the risks to human health is minimal.
The extent to which PAHs from contaminated animal feeds are transferred
into meat for human consumption is not clear (Deutsche Forschungsgemein-
schaft, 1976). The range of concentration values agrees with those
measured by Hfitt et al. (1978) who studied corn contamination in
high-temperature drying systems using a variety of fossil fuels. The
PAH levels are of the same order of magnitude as those found by Joe
et al. (1982) for barley drying.

Wood chips and corncobs are similar biomass fuels with respect
to their effect on the PAH condensation/absorption by corn. There was
one exception. Corncobs resulted in a very small quantity of naphtha-
lene in the dried corn, wood chips did not.

Chakraborty and Long (1968) studied the formation of soot and
PAHs in flames. They concluded that high temperatures in the pyrolysis
zone favor the formation of carbonaceous residues and lead to a reduc-
tion of PAHs. Also, PAHs decrease with increasing oxygen concentra-
tion during pyrolysis, almost reaching zero at 20% excess oxygen.

They also found that the optimum temperatdre fer' PHA synthesis in the

pyrolysis zone is 700°C. At temperatures above 1000°C there is no

217

formation of PAHs. In the MSU biomass furnace the primary combustion
zone (where pyrolysis occurs) temperature varied from 870°C to
1100°C at steady-state due to the preheating of the primary combustion
air. This temperature range is above the optimum range for PAH forma-
tion. Thus, the low levels of PAH contamination.

Duplicate corn samples run on different days showed no
detectable levels of the following metals: arsenic, cadmium, chromium,
mercury, selenium, lead, and thalium. The dectection limits for these

metals are given in section 5.1.2.

6.8 General Observations
Some general observations on the CVCF/IBCF system are presented
in this section. The adequacy of design of the major system components

and maintenance requirements are discussed.

6.8.1 Fuel Feed System

The fuel feed system consists of a self-unloading wagon and a
feed hopper. The fuel feed system performed satisfactorily. Wood
chips and corncobs without husks flow very smoothly through the system.
The fuel hopper is coated with graphite-based paint. This feature, as
well as the vibrator attached to the hopper, facilitates the smooth
flow through the hopper. Corncobs with a large percentage of husks
do not flow as smoothly as wood chips and husk-free cobs. There is no
bridging of fuel at the bottom of the hopper.

Backfiring through the feed hopper is effectively prevented by
maintaining a slight vacuum in the combustion chamber. The vacuum is

maintained by the eductor system and the drying fan.

218

The fuel feed rate is manually controlled by varying the speed
of the -stoker-auger. The fuel feed rate is adjusted to maintain a
specified drying temperature. When properly set, the feed rate is
equal to the fuel combustion rate. There is no accumulation of
unburned fuel in the combustion chamber during steady-state furnace
operation. The drying air temperature fluctuates by i 5°C from the
set temperature. The fuel feed rate control can be automated to
respond to the drying air temperature. However, this increases the
cost of the furnace without a corresponding improvement in the
furnace operation.

The fuel feed system requires little maintenance.

6.8. 2 Grate Performance

The grate in the CVCF furnace requires periodic maintenance
to remove clinker from the grate surface. The ash receptacle should
be cleaned at least once every other day. Accumulation of ash is
about 1 to 2 kg/h depending on the fuel type and the fuel feed rate.
Most types of wood contain 0.1 to 2% ash, while corncobs contain
approximately 1.6% ash (see Table 2.2). Husklage (a mixture of
corncobs and husk) has a higher ash content (about 3%). The ashes
should be removed daily.

The formation of clinker is a serious problem in biomass
furnaces. The clinker blocks the grate openings and therefore, limits
the undergrate combustion air supply. Husklage produces more clinker
than wood chips and corncobs. The wood chips used during the testing

of the furnace produced relatively small amounts of clinker compared to

219

corncobs and husklage. The clinker should be removed at least once

a week if wood chips are burned and more often when husklage is used.
The firebricks randomly spaced on the grate were effective in

insulating the grate from the high temperature in the combustion cham-

ber. Also, the undergrate airflow cooled the grate from the underside.

Thus, the grate design is satisfactory. However, the grate required

replacement due to oxidation and other chemical reactions on the

grate.

6.8.3 The Combustion Chamber

The lower combustion chamber is lined with firebricks; the
upper chamber and the inner lid are covered with a blanket of insula-
tion with desirable refractory characteristics (the price of the
insulation is 25% firebrick). The firebricks maintain dimensional
integrity and desirable refractory and heat capacity properties without
deleterious effect from biomass fuel abrasion. The insulation and the
firebricks effectively protect the metal parts (the inner steel cylin-
der and the inner lid) from the high combustion flame temperature
(1,000-1,500°C) and the oxidizing atmosphere.

The maintenance of the combustion chamber includes replacement
of the tuyeres and bricks. It is not known how long these components

will last.

6.8.4 Furnace-to-Dryer Duct
The heat from the furnace is recovered by means of heated air

blown around the furnace and mixed with the flue gases. The hot air

220

is blown into the bin plenum by a centrifugal fan via a rectangular
duct.

Approximately 95% of the heat loss from the system is through
the duct linking the furnace to the dryer. The heat loss varies from
300 to 500 MJ/hr depending on the ambient and the drying air tempera-
tures. The heat loss from the duct can be reduced considerably by

insulating the duct and the bin plenum.

6.8.5 Safety

Moving components of the furnace are adequately covered to
prevent operator contact with injury. Carbon monoxide poisoning is a
potential hazard. Exposure to the drying air inside the drying bin
should be limited. Some sparks and a small quantity of fly ash reach
the plenum and may result in some occasional smoldering of the

collected corn dust.

6.8.6 Miscellaneous

Unprotected bin and furnace components are likely to undergo
accelerated corrosion due to exposure to the products of combustion.
The rate of deterioration of these components is unknown. After two
seasons of furnace operation, the corrosion and soot deposits in the

bin interior are minimal.

CHAPTER 7

SUMMARY

A direct combustion concentric vortex-cell biomass furnace has
been designed, built, and incorporated in an in-bin counterflow dryer
(IBCF). The furnace is 1.2 m in diameter and 3.7 m high and has a
maximum energy output of 690 kW. It is made of two concentric steel
cylinders. The inside cylinder is lined with firebricks for insula-
tion.

The furnace/dryer system was tested under different ambient
and drying conditions to determine the technical and economic feasi-
bility as an alternative system for grain drying. The experimental
results were compared with the performance of the propane fueled
IBCF dryer.

The results demonstrate that the concentric vortex-cell biomass
furnace (CVCF) is a technically feasible alternative to fossil fuel
burners for grain drying. The corn drying capacity of the experimental
biomass fueled system is 2.0-3.0 wet tons per hour depending upon the
drying air temperature, the grain depth, and upon the initial and the
final grain moisture content. The drying capacity of the biomass
fueled system is approximately 90% of the propane system. The reduction
in the drying capacity is a result of the reduction in the flow rate of

the drying air mass due to the heating of the air before the fan entrance.

221

222

The CVCF/IBCF system operates at about 70% efficiency in con-
verting biomass fuels into energy for grain drying. Most of the heat
losses (about 95% of the total) occur between the furnace and the dry-
ing bin plenum. The heat loss from the furnace is about 5% of the
total heat loss. This value was minimized by covering the top one-
third of the furnace (1.2 m) with a shroud which reduces the radiation
and convective heat loss from the furnace. Also, the lower half of the
furnace is located below the ground level, thereby further reducing
the low heat loss from the furnace.

The average net specific energy consumption (SECO, kJ/kg of
water removed) of the biomas fueled system is about 6,100 kJ/kg com-
pared to an average of approximately 4,600 kJ/kg for a propane fueled
system.

Approximately 140 metric wet tons of corn was dried during the
testing of the CVCF/IBCF drying system. The biomass furnace performed
satisfactorily. There is no visible discoloration of or objectionable
odor in the corn dried with the CVCF/IBCF system. Chemical analyses of
the corn did not show dangerous levels of polycyclic aromatic hydro-
carbons (PAHs) or heavy metal concentration; the levels detected were
of the same order of magnitude as found in corn dried with propane
or natural gas.

The CVCF furnace is dependable. Drying air temperature control
is stable (within i 5°C of the desirable temperature) and is maintained
by regulating the fuel feed rate. The vibrator installed on the feed
hopper prevents fuel bridging, and maintains a continuous flow of

biomass fuel to the combustion chamber.

223

The grate system requires periodic maintenance to remove
clinker, especially after burning corncobs. The period between
cleanings depends upon the fuel feed rate and upon the ash content of
the fuel. 0n the average, the grate should be cleaned at least once
a week.

The ash accumulation in the receptable is on the order of 1 to
2 kg/h depending on the type of fuel used and the fuel feed rate.

This is due to the low ash content (0.2 to 3 percent) of wood and
corncobs.

The energy drying costs (exclusive of labor and fixed costs)
are 30 to 40% of the energy costs of a propane system. However, the
total break-even costs of the biomass system (including energy costs,
labor, maintenance, depreciation, and taxes) are about 20% higher than
of a propane system (at 1983 prices). The high costs are attributable
to the high capital investment (approximately $16,000) required to
convert a propane system into a biomass system. If the propane costs
increase by about 70%, the biomass system becomes competitive with the
propane fueled system. The concentric vortex-cell biomass furnace
does not have an economic advantage as a grain dryer heat source at

the current fossil fuel prices and interest rates.

CHAPTER 8

CONCLUSIONS

The main conclusions of this research are:

1. The concentric vortex-cell biomass furnace (CVCF) is a
technically viable alternative to propane and natural gas burners for
grain drying.

2. The CVCF furnace requires slightly more operator atten-
tion than a propane system.

3. The drying air temperature is stable during the steady-
state Operation of the furnace, and can be maintained at 70 to 95°C
by regulating the fuel feed rate.

4. The drying capacity of the CVCF coupled to an in-bin
counterflow (IBCF) grain dryer is about 90% of a propane fueled
IBCF system operating at the same drying conditions.

5. The average net specific energy consumption of a CVCF/IBCF
systemisabout 30% higher than of a propane fueled system. This is
due to the reduced capacity and additional heat loss in a biomass
system.

6. The energy costs (exclusive of labor and fixed costs) of
the CVCF/IBCF system are 30 to 40% lower than the energy costs of a
propane fueled IBCF system. The total break-even cost of operating a

biomass system is 20% higher than of the propane system, due to the

224

225

high investment of converting a propane fueled system into a biomass
fueled system.

7. The CVCF furnace does not have immediate economic feasi-
bility at the current propane prices and interest rates. The propane
prices have to increase by about 70% for the biomass system to become
competitive with propane systems.

8. The corn dried by the CVCF/IBCF system has no visible dis-
coloration or objectionable odor; and does not contain dangerous levels
of carcinogenic polycyclic aromatic hydrocarbons or heavy metals.

9. The CVCF furnace model developed in this thesis accurately
predicts the required biomass fuel feed rate and heat loss at differ-
ent drying airflow rates, drying air temperatures, ambient temperatures,
and fuel moisture contents. The model can be modified to simulate the
performance of other furnace types.

10. The in-bin counterflow grain drying model presented simu-
lates the intermittent action of the IBCF system. The values of the

simulated parameters are close to the experimental values.

CHAPTER 9

SUGGESTIONS FOR FURTHER RESEARCH

The following recommendations are suggested as topics for

further research:

1.

Design of automatic furnace shut down in the event of:

a. drying bin empty,

0'

drying bin failure,

c. extremely high or low drying air temperature,

d. accidental fire.

Determination of the optimum dimensions and configura-
tion of biomass furnace eductors.

Determination of the economic feasibility of insulat-
ing the CVCF furnace, the duct work between the
furnace and the drying bin, and the drying bin plenum.
Investigation of alternative insulating materials for
the combustion chamber of a concentric vortex-cell
biomass furnace (CVCF).

Determination of the effect of tuyere size and tuyere
angle on particulate emission.

Testing of other biomass materials as fuel, such as

coal, apple pomice, rice hulls, etc.

226

10.

11.

227

Application of CVCF furnace to other dryer types.
Adaptation of the furnace model to simulate other
biomass furnaces.

Dimensional analysis and similitude modeling of the
CVCF and the IBCF systems.

Statistical analysis to determine the variations in

the experimental and the simulated values of fuel feed
rate, heat loss, grain moisture content, and grain dry-
ing capacity of the CVCF/IBCF system.

The effect of precipitation and wind on the heat loss

from the CVCF/IBCF drying system.

APPENDICES

228

APPENDIX A

SAMPLE 0F CVCF COMPUTER INPUT/OUTPUT DATA

229

 

Sample of CVCF Computer Input/Output Data.

FUEL TYPE: CCRNCUES=1,MCCD=8.E. HUUDCHIFS
FUEL FEED RATE,RGKHR:188. 188 88
FUEL iEIETEEE, PERCENTCNEFEE. 28.68
AMBIENT TEMFEEETEEE, C:4. 4.88
PERCENT EELETIEE HUMIDITY:7S. 75.88
DRYING EEEEEEEE, EEEI. 1.:8
EezEe- EIF, PERCENT:EE EB.88
COMBUSTIDN EIF. 1F’H: 757 7?
AIR:FUEL EETIE- 7 58
T EM F‘ E RHT Ll F! E '5‘ :
AMBIENT: 4.88:
FEEHEETEE EEME. EIE: 123.9EE
FEEHEEE FLAME: 1243.4EE
FEEEILETIEN FLUE GEE: 783.54C
EILETEE FLUE GEE: 181.87C
DFrIvS (BIN PLENUM}: 7?.62C
ULzTER FEENEEE EELL: 14E.EEE
MEEE FLEEE:
FLEL FEEE KG/H: 188.88
FEEL MEIETEEE, we: 28.88

FURNACE AIR SUPPLY, RGJHG-FUEL: 14.43
CMM: 19.32

EWCESS AIR, PERCENT: 50.88
PRECILI jTIDN FLUE GQS, RBEH: 1547.85
CNN: 77.14
DILUTED FLUE GAS, KGKH: 16388.67
CNN: 28?. 02
DRYING AIR, CMM: 274.11
DRYING HIR HE‘EDLUT E HLIV CITY: .8876
TOTéL HEAT OUTPUT, MJEH: 167 7.20
TCTQL HEAT LOSS, MJ./H: 44S. 95
BIN PLENUM HEAT LOSS, MJKH: 58.24
FURNACE EFFICIENCY PERCENT: 95.38
SYSTEM EFFICIENCY, PERCENT: 73.23
ASH ACCUMULATIUN, KG/H: .94
UNBURNED FUEL ACCUMULATION, RG/H: .88

230

APPENDIX B

EXPERIMENTAL RESULTS

231

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APPENDIX C

CVCF COMPUTER PROGRAM

239

C.
C*
C‘I'
c.
C*
C*
c.
c*
c.
C*
c*
CI‘
c.
Cfl’
C*
C*
C.
C*
c*
C.
C.
C.
Ci
c*
c.
C*
C'I'
c*
C.
C*
C.
C*

CVCF COMPUTER PROGRAM

PROGRAM BURNER(OUTPUT,INPUT)

THIS PROGRAM COMPUTES THE HEAT LOSSES FROM THE MSU CONCENTRIC
VORTEX-CELL BIOMASS FURNACE. IT ALSO COMPUTES THE TEMPERATURES
AND THE MASS FLOWS IN THE VARIOUS SECTIONS OF THE FURNACE.

THE PROGRAM CONSISTS OF THE MAIN PROGRAM 'BURNER', SUBROUTINES
'PROPATY', 'AIRPTS', 'FLUPTY', AND THE FUNCTIONS 'RUTA',
'SIMPSON', ENTHAL, 'TABLE', HEATF', 'TADI', AND 'GMOW'.

THE SUBROUTINES FUNCTIONS ARE AS FOLLOWS:

PROPATY: CONTAINS TABLES OF THERMODYNAMIC PROPERTIES OF
THE MAJOR GASES INVOLVED IN COMBUSTION.

AIRPTS: COMPUTES THE PROPERTIES OF AIR AT A GIVEN TEMPERATURE
CALLS THE FUNCTION TABLE WHICH INTERPORATES THE DATA
CONTAINED IN THE SUBROUTINE PROPATY.

FLUPTY: COMPUTES THE DYNAMIC VISCOSITY AND THERMAL
CONDUCTIVITY OF FLUE GASES.

RUTA: SOLVES FOR THE JACKET INNER.WALL TEMPERATURE BY
DOING THE RADIATIVE, CONVECTIVE, AND CONDUCTIVE HEAT
BALANCE.

SIMPSON: COMPUTES THE ENTHALPY CHANGE OF FLUE GASES AT A
GIVEN TEMPERATURE BY THE SIMPSON INTERGRATION OF
THE SPECIFIC HEAT FUNCTIONS IN THE SUBROUTINE ENTHAL

ENTHAL: CONTAINS SPECIFIC HEAT EQUATIONS FOR GASES.

TABLE: INTERPORATES THE TABULATED DATA.

HEATF: DETERMINES THE HEAT REQUIRED TO HEAT THE AIR AND THE
FUEL TO THE COMBUSTION TEMPERATURE ABOVE 25C.

TADI: DETERMINES THE ADIABATIC FLAME TEMPERATURE GIVEN THE
FUEL HEATING VALUE, THE INITIAL FUEL AND THE AIR
TEMPERATURES, AND THE PRODUCTS OF COMBUSTION.

GMOW: COMPUTES THE MOLECULAR WEIGHT, SPECIFIC HEAT, AND THE
TOTAL MASS OF FLUE GAS.

COMMON/FRACTNS/COM(20),GAS(20),WMOL(20),CP(20),CM2,CM3,CM4,CMO
COMMON/PROPETY/AK(30),ACP(30),APR(30),ADEN(30),AV(30),CFMF(10)
COMMON/FLUE/DVCOZ(20),CKCOZ(20),DVCO(20),CKCO(ZO),DVH20(20),
+CKH20(20),DVH2(20),CRH2(20),DVN2(20),CKN2(20),DVOZ(20),CKOZ(20)
COMMON/SMS/SMLA,SMLFLU,DIFA,DIFLU,KA,KFLU,SMST,DIFST,KST,COMED(10)
cmn/mssnam

DIMENSION NAME(20),A(20),R(30),X(50),TFUR(50),VISH(20):QL(20)
DIMENSION GASHI(10)

240

C.

c.
C.
C.

c.
C.
C.
C.

33
43

c.
C.
c.

80

241

EXTERNAL HEATF,FLUPTY,TABLE,PROPATY,RUTA

DATA PATM/14.696/

DATA C,H,OXY,CN,SUL,ASH/46.83,5.9S,45.025,.42,0.,1.6/

DATA FUEL,XMFW,TAMB,RH,PFF,BINP/90.,12.l,5.79,70,2.125,3.44/
THE A'S ARE THE AREAS OF THE VARIOUS SECTIONS OF THE FURNACE.
DATA (A(L),L=l,6)/.732,10.167,1.167,14.009,2.627,14.009/
THE R'S ARE THE THERMAL RESISTANCES ACROSS VARIOUS SECTIONS OF
THE FURNACE INCLUDING CONVECTIVE, RADIATIVE AND CONDUCTIVE
THEMAL RESISTANCES.

DATA (R(M),M=1,18)/.527,15.814,.978,S.,.312,2.385,.286,
+7.88,9.362,4.029,.125,.125,3.75,15.228,1.0075,.157,.15,.008/
DATA AAO,AA2,AA3,AA4,AP/.09931,.00456,.0324,.0243,1.459/
DATA (WMOL(K),K=l,B)/44.01,28.0l,28.016,46.,30.,18.,32.,2./
TRK(R)=R/l.8

THR(C)=1.8*C

TKF(Z)=1.8*Z-459.67

TFK(Z)'(Z+459.67)/1.8

CALL PROPATY<TAMB)
THE COMS ARE THE COMPONETS OF THE FLUE GASES CONSIDERING THE
COMBUSTION AIR AND FUEL AS THE REACTANTS.
THE COMEDS ARE THE COMPONENTS OF FLUE GASES AFTER DILUTION WITH
EDUCTOR AIR.

NAME(1)'9H CORNCOBS

NAME(2)'9HWOODCHIPS

DO 33 N31,10

COM(N)=COMED(N)=O.

CONTINUE

PRINT 133

READ 329,TYPE

IF(TYPE .EQ. l.)PRINT 83,NAME(1)

IF(TYPE .EQ. 0.)PRINT 83,NAME(2)

TJAK=TPRE84OO.

WHAT'TYPE

IF(TYPE .GE. l.)GOTO 80
C, OXY, CN, SUL, ASH, AND H ARE THE PERCENTAGE COMPONENTS OF
FUEL BEING CARBON, OXYGEN, NITROGEN, SULFUR, ASH AND HYDROGEN,
RESPECTIVELY, AS GIVEN IN THE FUEL ULTIMATE ANALYSIS.

C=52.02

OXY'40.61

CN=.O9

SUL=.Ol

ASH81.18

H86.09

CONTINUE

PRINT 4

READ 329,FUEL

PRINT 200,FUEL

PRINT 8

READ 329,!MFW

PRINT 200,3MFW

XMDBIXMFW/(IOO-XMFW)

PRINT 5

c.
c.
C.
C.
C.

C.
c.
c.

242

READ 329,TAMB
PRINT 200,TAMB
PRINT 99
READ 329,RH
PRINT 200,RH
PRINT 6
READ 329,3INP
PRINT 200,BINP
BINP=BINP/.2482342855
NAMEéthWIND
NAMES=BHVELOCITY
NAMEGc3HM/s'
PRINT*,NAME4,NAMES,NAME6
READ 329,VWIND
VWINDBO.
PRINT 200,VWIND
THE PPP IS THE COMBUSTION AIR PRESSURE AS A FUNCTION OF THE DRYING
PLENUM AIR PRESSURE (BINP).
PPP=.3075*BINP+I.232
IP(BINP .LE. 0.)PPP=2.25
NAME3=5HR(5)
TAMBsTAMB+273.IS

C* FCFM IS THE TOTAL CFM OF AIR INTO THE FURNACE
C* FCFM INCLUDES THE EDUCTOR AND THE COMBUSTION AIR,
C* AND IS DETERMINED FROM THE CHARACTERISTIC FAN EQUATION.

FCFM=TABLE(CFMF,SMST,DIFST,KST,PFF)
W=HADBRH(TKR(TAMB),RH/lOO.)
SVO=VSDBHA(TKR(TAMB),W)
CMO=FCFM/SVO*.4S36

TEK=TAMB+4.3
CMOBCMO*(-.05*BINP+1.19)

PRINT 111

READ 329, EXCS

PRINT 200,EHCS

CMOML‘CMO/28.97

C* CFM3 IS THE AIRFLOW AROUND THE FURNACE, (CFM)
C’ CM3 IS THE AIR MASS FLOW AROUND FURNACE (KG/MIN)

cuv3=(479.93-10.65*BINP)
CFM3=CMV3*15.708
CM3=CEM3/SVO*.4536
cuv3=CMv3*.3048

AAsC/12.01115

BB-H/1.00797

CC-CN/14.0067

nn=SUL/32.064

EE=OXY/15.9994
STOXY=AA+BB/4.+CC+DD-EE/2.
FUEL=PUELI60.
FUELDa-PUELn-FUEL*(1.-xurw/Ioo.)
AMPKGISTOXY‘4.76/IOO.
AMOL$=5T0X!*4.76*ruan/Ioo.
AIRMoLs-Auons*(zxcs+loo.)lloo.

 

C.
c.

c.
.c*
C.

11

39

C.
c.
c.
C.
C.

243

IP<AIRMOLS .LT. AMOLS)PUELDE=AIRMOLS/AMPRG

IP(AIRMOLS .LT. AMOLS)AMOLS=AIRMOLS

CM2=AIRMDLS*28.97

CMC=CM2*60.

CMR=CMC/(FUEL*60.)

FUELNB=FUEL*60.-FUELDB*(l.+XMDB)*60.

PRINT 899

PRINT 200,CMC

PRINT 900

PRINT 200,CMR

CMED=CMO-CM2

ExT1=AIRMOLs-AMOLS
HHV Is THE PUEL HIGHER HEATING VALUE, NJ/NG As A PUNCTION OF

- THE PUEL COMPOSITION (ULTIMATE ANALYSIS).

ENV=18650.

IP(TYPE .EQ. o.)EEVs20965.

HHVT=HHV*FUELDB

ENT2=CMOML-AMOLS

STOCRA=AMOLS*28.97/PUELD

ACTRA=CM2/PUELD

EORATIO=STOCRA/ACTRA

COM(I)=AA*PUELD/IOO.

COM(3)=.7899*AIRMOLS+CC/2*PUELD/Ioo.
COM(6)=NMDE*PUELD/Ia.015+I.608*AIRMOLS*N+SE/2.*PUELD/IOO.
COM(7)=.2101*ExTI
DNA IS THE TOTAL MOLES IN THE PLUE GAS.
GAS<I> Is THE PRACTIONAL MOLECULAR COMPOSITION OP COMPONENT I
IN TEE PLUE GAS.

QNA=COM(1)+COM(3)+COM(6)+COM(7)

DO 11 NsI,7

IP(COM(N) .EQ. o.)GOTO 11

GAS(N)=COM(N)/QNA

CONTINUE

COMED(1)-COM(1)

COMED<3>=.7899*CMOML+CC/2*PUELD/Ioo.
COMED(6)-xMDa*PUELD/Ia.OIS+I.6081*CMOML*N+EE/2.*PUELD/Ioo.
COMED(7)-.2101*ExT2
ONA2=COMED<I)+COMED<3>+COMED<6)+COMED(7)

DO 39 N=I,7

IP(COMED(N) .Eo. o.)GOTO 39

GASED<N>=COMED<N>/ONA2

CONTINUE

AIRMAss-CMz

KCYCLE=1

TADR IS THE PIRST ESTIMATE OP ADIAEATIC PLAME TEMPERATURE
AS GIVEN BY TILMAN ET AL. (1981).

TAD2 IS THE ADIAEATIC PLAME TEMPERATURE TAKING INTO ACCOUNT
THE REACTANTS INITIAL TEMPERATURE AND THE HEAT REQUIRED POR
PRENEATING THE REACTANTS (FUEL AND AIR).
TADR=1920.-2.*xMDE*IOO.-s.2*AES(ExCS)

TADEs-TADN

1300 CONTINUE

C.
C.
C.

15

244

IF(KCYCLE .EQ. I)CPA=TAELE(ACP,SMLA,DIPA,NA,TPRE)
CON3=NEATP<TPRE,CM2,PUEL,xMDE,TAME)
TAD2=TADI(COM,TADES,ONA,WMI,CMG,CPG,CON3,NEVT,ENP)
TAD2=<TAD2+TADK)/2.

TPRE=TJAN

IF(KCYCLE .EQ. 1)TP-TAD2-CM2*CPA*(TPRE-TAMB)/<CMG*CPG)
IP(TP .LE. TJAK)PRINT 93

IP(TP .LE. TJAK)GOTO 91

TADESeTADZ

E=.2

GMsGMON<COM,TP,CPG,CMG,ONA)

AWS=7.7587

ANT=.456

ANIsAWS+ANT

AGR-.5

AP=4.4*AGR
EC02820-483./TP*(.424*(GAS(6)+GAS(1))*Ioo.)**.45
ESOOT=1.-EXP(-.6*E*SORT(AGR))
EPS=ESOOT+ECOZHZO-(ESOOT*ECOZHZO)
R(19)=(1.-EPS)/(AP*EPS)+I./(AP*.75)+.0612
STEETz=s.669E-8

SE=STEETz

RU=8.2E—3

REOG=WMI/(RU*TP)

CC1=EPS*STEETz*4.*AGR*60.
VEG=CMG/(RNOG*AGR)*.65/60.

AN2L=2.027E-3

TNZL=(TF*2.+TPRE)/3.
AIRD=TAELE<ADEN,SMLA,DIPA,NA,TN2L)
VENZL=CM2/60.*.35/(AN2L*AIRD)
VEMAxssORT(VEG**2.+VENzL**2.)

COMPUTE VISCOSITY 8 THERMAL CONDUCTIVITY OF FLUE GASES

CONRA-o.

CONDMIx-UMIxso.

DO 27 I-I,6

IP(COM<I) .Eo. O.)GOTO 27
VISK(I)-FLUPTY(TF,I,1)

RATIo-I.

DO 15 J-I,6

IP(COM(J) .EQ. o.)GOTO Is

IP(J .Eo. I)GOTO Is
VISK(J)-FLUPTY(TP,J,1)
P1-1.+SQRT(VISK(I)/VISK(J))*(WMDL(J)/WMDL(I))**.25
PI-P1**2.
P2-2.*SORT(2.)*SORT(I.+NMOL(I)/wMOL(J))
PEI-PI/Pz
RATIOsRATIO+PNI*(GAS(J)/GAS(I)>
CONRAsGAS(J)/Ioo.*PNI+CONRA

CONTINUE

COND-PLUPTT(TP,I,2)

245

CONDMINsCONDMIN+<GAS<I)/IOO.*COND)/CONRA
UMIX=VISK(I)/RATIO+UMIX

27 CONTINUE
PRG=CPG*IOOO.*UMIx/CONDMIx
HI=VEMAx*RnOG*CPG*IOOO./PRG**(2./3.)*.0125
CC2=EI*4.*AGR*60.
CP-CCZ/GO.
OPUNCsCMG*CPG*IOOO./60.*(TAD2-TP)
RIT=O
TSXY=.0552*TAMB**1.S
EES=SE*TSKT**4
TPPsTxrcTP)
TP2=.6942*TPP-65.6386
T2=TPN<TP2>
EEZeSE*T2**4
TP7=.2449*TPP-259.41
T7=TE7=TPK<TP7>
TsaCMG*CPG*TP+CMED*TEN
T8=T8/(CMG*CPG+CMED)
EBB=SB*T8**4

10 CONTINUE
TPM=<T7+TAMB)/2.
CALL AIRPTS<TPM,DEN,SKS,VIS,PRN,CPS)
PETA-1./TPM
GR-(EETA*9.8*(T7-TAME)*1.2192**3)/VIS**2
RA-GR*PRN
Ev1-4.
Ev2-Ev1
IP(vaND .GE. 3.)Ev1-1.¢2
IF(VWIND .GE. 3.)Ev2=.9s
HA78EV1*((T7-TAMB)/2.4384)**.25
IP(RA .Gr. 1E+9)HA7-Ev2*(T7-TAMB)**(1./3.)
OCOI=1¢.*EA7*(T7-TAME)

C* PORCED CONVECTION
RE-VWIND*1.2192/VIS
NNU=.43+.5*RE**.5*PRN**.38
IP<RE .GT. IE+3)HNU=.25*RE**.6*PRN**.38
OCOP-4.67*SNS/1.2192*NNU*(T7-TAME)
IP(vaND .LT. 3.)OCOP-O.
OCO7-QC07+QCOP
TJAP=.1814*TPP-156.6478
TJAN-TPK(TJAP)
CALL AIRPTS<TJAR,DEN,SNS,VIS,PRN,CPS)
vsv-CM2/60./(DEN*.2137>
IP(NIT .EO. O)T67¢T7
TDI-T67-TAME
TDI-AES(TD1)
TD2-T57-TJAR
TD2=AES<TD2>
TD3-TD1-TD2
TDB-ABS(TD3)
TLM-TD3/ALOG(TD1/TD2)

20

23

1401

246

DE*.254

CALL AIRPTS(TLM,DEN,SRS,VIS,PRN,CPS)
H67=SKS*.023/DE*(DE*V67/VIS)**.8*PRN**(l./3.)
QG7=12.842*H67*(T7-TLM)

EBA=SB*TAMB**4

QS7=QCO7+Q67-EBA/R(17)

QC6=Q67*1.5

IF(KIT .EQ. 1)QC6=10.16667*H67*(T6-TLM)
Q567=QC6+QS7*R(l7)/(R(16)+R(l7))
RSl=R(l7)/(R(16)*(R(16)+R(l7)))-1./R(16)
IF(KIT .EQ. 0)T6=T7*1.5

K82

X(1)=T6*.925

X(2)=T6*l.5

CONTINUE

38K+1
FX1=RUTA(QSG7,RSI,CF,TF,R,X(H-2))
FX2=RUTA(QSG7,RSl,CF,TF,R,X(K-l))
X(K)=(FX2*X(K-2)-FX1*X(K-l)>/(FX2-FX1)
ERROR=ABS(X(K)-X(K-l))/2.

IF(ERROR .LE. 1E-10)GOTO 23

GOTO 20

CONTINUE

T68ABS(X(K))

KITSKIT+1

EBG=SB*T6**4
TW1=(Q567-EBG*RSI+T6/R(18))*R(18)
EB7=T7**4.*SB

IF(T7 .LT. TE7)T7=TE7

T67'(T6+T7)/2.

IF(KIT .EQ. 1)GOTO 10
QR7=.75*(EB7-EBA)/R(l7)+.25*(EB7-EBS)/R(l7)
TF7-THF(T7)

AA4=1.1675

TF481.0605*TF7+15.3076
TAF=.OB*TKF(T8)+26.

V4'2.413

T4-TFK(TF4)

EBGISB*T4**4

TA-TFMTAF)

DEI.61

TW‘(T4+TAMB)/2.

TDIITO-TAMB

TDlIABS(TD1)

KEYl=0

CONTINUE

TDZ-T4-TA

TDZSABS(TD2)
DTM=(TDl-TD2)/ALOG(TD1/TD2)

CALL AIRPTS<DTM,DEN,SNS,VIS,PRN,CPS)
H4ISKS/DE'.023*(V4*DE/VIS)**.8*PRN**(1./3.)
Q40!AA4*HC*(T4-DTM)

247

TIN4=<T4+TAME)/2.
TDl-T4-TAMB
TDI=AES<TDI>
TD2=T4-TIN4
TD2=AES<TD2>
DTM=(TDl-TD2)/ALOG(TDl/TD2)
CALL AIRPTS<DTM,DEN,SNS,VIS,PRN,CPS)
CA4=I.069
VIN4=CM2/(DEN*CA4)/60.
EIN4=SNS/.8636*.023*(.8636*VIN4/VIS)**.8*PRN**(1./3.)

' QIN4=AA4*NIN4*(T4-DTM)
DJ4=(Q40+QIN4)*R(7)+SB*T4**4
EEIsSE*Tw1**4
EEII=SE*(TwI*.65)**4
RTle/R(1)+1/R(2)+1/R(4)+1/R(IO)
RT2=1/R(2)+1/R(3)+1/R(8)+1/R(9)
RT3=1/R( 4)+1/R( 5)+1/R(6)+1/R( 9)
RT4=I/R(6>+1/R(7)+1/R(8)+1/R(10)
RT5=1/<R(2)**2*RTI)-RT2
RT6=1/(RT1*R( 4)*R(2) )+1/R( 9)
RT7-1/(RT1*R(IO)*R(2>)+1/R<8)
RT8=1/(RT1*R(10) *R( 4) )+1/R(6)
RT9=RT8-RT6*RT7/RT5
DRIzEEII/R(I)+D34/R(IO>
DN2=-(E32/R(3)+DJ4/R(8)>
DN3=DJ4*RT4-EE4/R(7)
DR4=DN2-DxL/(RTI*R(2))
DR5-DN3-DN1/(RTI*R(IO))
DJ3=(DK5-DK4*RT7/RTS)/RT9
DJZ=(DK4-DJ3*RTG)/RTS
DJI=<DK1+DJ3/R(4)+DJZ/R(2))/RT1
E33-AES(D33*RT3-(DJI/R(4)+DJZ/R(9)+DJ4/R(6)))*R(5)
T3-(E33/SE)**.25
T13=(T3+TW1*.65)/2.
EETA=1./T13
CALL AIRPTS<T13,DEN,sxs,VIs,PRN,CPS>
GR:.2775*EETA*AES(Ts-T13>/VIS**2

C* EOUTIONS 8-74 TO 8-76 GEEEART POR PREE CONVECTION
C* BETWEEN HORIZONTAL PLATES, BOTTOM NOTTER.

HNU=.068*GR**(1./3.)
IF(GR .LT. IE+5>HNU=.I95*GR**.25
IP(GR .LT. IE+4)NNU=I.
HSIHNU*SKS/ . 3048
AA3-1.0681
913=AA3*1.5*H3*(T3-T13)
CALL AIRPTS<TA,DEN,SNS,VIS,PRN,CPS)
DT-1.0973
VMT=.25*CM3
AT-I.806
VET=VMT/(AT*DEN)/60.
RE-VET*DT/VIS
ET-SNS/DT*.664*PRN**(1./3.)*RE**.5

45

47

248

TAP=(TAME+TA)/2.
QT3-AA3*NT*(T3-TAP)
DJS=(QIB+QT3-DJ3/R(5)+E83*(1. /R(5)+1. /R(5)))*R(5)
RT10=1. /R(5)+1. /R<11)+1. /R(13)
RT11=1. /R(11)+1. /R(12)+1. /R(14)
RT12=1./R(13)+1./R(14)+1./R(15>
RT13-R(11)/(R<13)*R(14))+RT12
DK6=DJS*RT10-EBB/R(5)
DK7=-(Ene/R(15)+DJ5/R(13))
DN8=DN7-DN6*R(11)/R(14)
DJ7s-Dx8/RT13
DJ6=(DK6-DJ7/R(13))*R(11)
EBS=ABS(DJ6*RTll-DJ5/R(11)-DJ7/R(14))*R(12)
Ts=(EES/SE)**.25

T5=(TA*5+T5)/6.

IP(TS .LE. TAME)T5=TA
0T03=AT*NT*(T5-TAP)
TA-((OT3+QTD3)*.06+CM3*CPS*TAME)/(CM3*CPS)
KEY1=KEY1+1

IF(KEYl .LT. 2)GOTO 1401
TP5=TKF(T5)

AT-2.6268

ORTsSE*AT*(T5**4—T5RT**4)
TTP=(T5+TAME)/2.

CALL AIRPTS(TTF,DEN,SKS,VIS,PRN,CPS)
IF(VWIND .LT. 3.)GOTO 45
RE-VWIND*1.6/VIS
HT5=SKS/l.6*.5*RE**.5*PRN**(1./3.)
GOTO 47

CONTINUE

EETA=1./TTP
GR-43.6955*EETA*(T5—TAME)/VIS**2
RA-PRN*GR

Ev3=7.

ETSsSNS/1.6459*EV3*RA**.2

CONTINUE

X(KCYCLE)-TF

KCYCLE=KCYCLE+1
OCTuAT*ETs*(T5-TAME)
QLT=QR7+QRT+QCT+OCO7+QT3+OTD3
QLN=(OR7+QRT+OCT)*3.6

IP(NCICLE .GE. 10)GOTO 1310

CALL AIRPTS(TA,DEN,SKS,VIS,PRN,CPS)
QLPT=QLT+CM2*CPS*(TJAN-TAME)*60.
QEVApstDE*PUELD*1.353E+s
QLFT=QLPT+QEVAP
TP-(CPG*CMC*60.*TAnz-OLPT)/(60.*CPG*CMG)
x(NCTCLE)-TP
ERR-ASS(x(xCICLE)-x(NCICLE-1))/2.
IP(ERR .LE. .05>GOTO 1310

GOTO 1300

1310 CONTINUE

C.
c.

c.
C.

C.
C.

c.

c.

439

433

c.

249

QINPUT=HHV*FUELD*60.
FUEP=<1.-QLN/QINPUT)*100.
GM=GMOW<COMED,T3,CPG2,CMG2,ONA2)

CPG2 IS THE SPECIFIC HEAT OF PLUE GASES AFTER EDUCTOR

CMGZ Is THE MASS OP THE FLUE GASES.
CALL AIRPTS<TA,DEN,SNS,VIS,PRN,CPS)
TxT=CM62*CPG2*T8+CM3*CPS*TA+Qc07*.06
TXT=TXT/(CMGZ*CPGZ+CM3*CPS)

TXT Is THE TEMPERATURE OF DILLUTED GASES

IMMEDIATELY AFTER THE FURNACE.

CMOL3=CM3/28.97
COM(3)=COMED(3)+.7899*CMDL3
COM(6)=COMED(6)+1.6081*CMOL3*W
COM<7>=COMED<7>+.2101*CMOL3
CMD=CPD=0.
QMD,CPD,CMD ARE THE TOTAL MOLES, SPECIFIC HEAT A MASS
OF THE DILLUTED PLUE GASES. '
OMD=COM<1)+COM(3)+COM(6)+COM(7)
GM=GMOW<COM,TxT,CPD,CMD,OMD)
BIF=COM(6)*WMOL(6)/CMD
COMPUTE HEAT Loss BETWEEN FURNACE a DRYING BIN
BIND=5.49
AREAB=12.7S
KCYCLE=0
IF(VWIND .LT. 3.>GOTO 1129
DUCT HEAT LOSS
TH=TXT
TFILM=<TNT+TAMB)/2.
CONTINUE
CALL AIRPTS(TFILM,DENA,SESA,VISA,PRNA,CPSA)
RED=vaND*.65/VISA
CTY=.66
IP(RED .GT. 5.7SE+4)CTY=.91*RED**(-.02445)
cx=19.
IP(RED .LT. 5.7SE+4)CX=15.
CRED=cx*RED**(-.28895)
KCYCLE-KCYCLE+1
NA=SRSA*CRED*RED**CTI*PRNA**(1./3.)
EA-EA/.65
ODUCT-zz.58*NA*(TU—TAMB)*.06
TTBN=TXT-ODUCT/(CPD*CMD)
TH=(TI'BN+TXT)/2.
TFILM=<TN+TAMB)/2.
IF(KCYCLE .LT. 2)GOTO 439
NCTCLan
TAVEuTTBN
TPILM-(mN+TAMB)/2.
CONTINUE
CALL AIRPTS(TFILM,DENA,SKSA,VISA,PRNA,CPSA)
REBN=VWIND*BIND/VISA
IP(REBN .GT. 1E+3)GOTO 1000
EQUATION 6-16, HDLMAN

250

UNU=<.43+.5*REEN**.5)*FRNA**.38
GOTO 1001

C* EQUATION 6-17, MOLMAN

1000 HNU=.125*REEN**.6*FRNA**.38

1001 HBIN=HNU*SKSA/BIND
QEIN=EEIN*AREAE*(TAVE-TAME)*.06
KCYCLE=KCYCLE+1
TEP=TTEN-QEIN/(CFO*CMD)
TAVE=<TEF+TTEN)/2.
TFILM=<TAVE+TAMB)/2.
IP(KCYCLE .LT. 2)GOT0 433
GOTO 1100

1129 CONTINUE
NCYCLE-o

C* FREE CONVECTION HEAT LOSSES, vaano.
TW=TXT
TFM=<TW+TAMB)/2.

533 CONTINUE
NCYCLE-NCYCLE+1
CALL A1RPTS<TFM,OEN,SNS,VIS,FRN,CFS)
EETA=1./TFM
GR=9.8*EETA*(Tw-TAME)/VIS**2.*.65**3.
GRERaGR*FRN
EVI-Ev2=5.25

C* EQUATION 7-25, HOLMAN
HNU=Ev1*GRPR**.25
IP(GRPR .GT. 1.E+13)HNU=EV2*GRFR**(1./3.)
QOUCT=22.58*HNU*SRS*(Tw-TAMB)*.0923
TTEN=TxT-QOUCT/(CP0*CMD)
Twr(TTEN+TxT)/2.
m=('rW+'rAMB)/2.
IP(KCYCLE .LT. 2)GOT0 s33
NCYCLEso
THE-TTEN
TFMs<TWE+TAME)/2

633 CONTINUE
NCICLE-RCYCLE+1
EETA-I./TFM
CALL AIRPTS<TFM,DEN,sxs,VIS,FRN,CFS)
EINL=.9138
GR=9.8*EETA*EINL**3*(TNE-TAME)/vIs**2
GRPR-GR*FRN
Ev1-Ev2-GRFR*3.E-9+1.25

C* EQUATION 7-25, UOLMAN
NNUsEv1*GRFR**.2s
IF(GRFR .GT. 1.E+13)HNU=Ev2*GRFR**(1./3.)
QEIN-SNS*HNU/EIND*AREAE*(TNE-TAMB)*.06
TEFsTTEN-QEIN/(CMD*CFD)
TWB=(TTBN+TBP)/2.
TFM=<TNE+TAMB)/2.
IP(KCYCLB .LT. 2)GOTO 633

1100 CONTINUE

 

C?

251

CALL AIRPTS<TBP,DEN,SKS,VIS,PRN,CPS)
DAIRv=CMD/DEN
CUFTAsOAIRv*35.3134/254.469
CFMT=CUFTA*254.469
QLH=QLH+(QBIN+QDUCT)*60.
SYEF=(l.-QLH/QINPUT)*100.
TEIN=TKF<TEP>

COMPUTE DATA FOR PRINTING

HHVW=QINPUT/FUEL/1000.
HHVD=HHV/IOOO.
ASHF=ASH/100.
NKGLE=2.204622
ASN=ASH*FUELD*60./100.
ASHLE=ASH*WKGLE
TOC=TAME—273.15
TOF=TKF<TAMB>
TPC=TJAK-273.15
TFF=TKF<TJAK>
TFC=TF-273.15
TPP=TKF(TF)
TXTC=TXT-273.15
=TKP(TXT)
TEDC=T8-273.15
TEDF=TKP(T8)
TENC=TEP-273.1s
TWC=T7~273.15
TWPeTKF(T7>
PRINT 150,TOC,TOF,TFC,TFF,TFC,TFF,TEDC,TEDF,TxTC,TXTF,

+TBNC,TBIN,TWC,TWF

FULB=FUEL*WKGLB*60.

VWIND=VWIND*3.6

QINPUT=QINPUT/1000.
WM?H8VWIND*5./8.

CMD'CMD/FUEL

CHMCBPCFM*.02837
PDPG=(CM2+CMED+FUEL*(1.-ASEP))*60.
PDLBBPDPG'WKGLB

CALL AIRPTS(T8,DEN,SKS,VIS,PRN,CPS)
CMMED-PDFG/DEN/GO.
CPMEDICMMED*35.3134
CMDLB=CMD*WKGLB*60.

FUEL‘FUEL*60.

CALL AIRPTSCTXT,DEN,SKS,VIS,PRN,CPS)
CHMDBCMD/DEN

CPMDSCMMD*35.3134

CMDBCMD'GO.

BTUM‘QINPUT'9.4783E-4

QLH‘QLH/IOOO.

QLBTUBQLB’9.4783E-l

PRINT 25,FUEL,PULB,XHFW,CHO,CMMC,FCFM,EXCS,PDPG,PDLB,

+CHHED,CPMED,CMD,CHDLB,CHHD,CFHD,DAIRV,CFHT,HIP

PRINT 37,QINPUT,BTUM,QLH,QLBTU,FUEP,SYEP,ASH,ASHLB

91

111
109

25
ca

37

1400

200
83
329
93
899
900

252

FNLB=FUELNB*NNGLB
IF(FUELNB .GT. 0.)PRINT 1400,FUELNB,FWLB
CONTINUE

PRINT 109

READ 329,TERM

IF(TERM .GE. 1.)GOT0 43

FORMAT(* *,*EXCESS AIR, PERCENT:*) -
FORMAT(*0*,/,* *,*CONTINUE? YES=1.; NO=0.*)

FORMAT(* *,*PERCENT RELATIVE HUMIDITY:*)

FORMAT(* *,*FUEL MOISTURE, PERCENT(NB)*)

FORMAT(* *,*AMBIENT TEMPERATURE, C:*)

FORMAT(* *,*FUEL FEED RATE,xG/HR:*)

FORMAT(* *,*DRYING PRESSURE, RPA*)

FORMAT(*1*,*FUEL TYPE: CORNCOBS=1,w00D-0.*)
FORMAT(*O*,/*O*,“TEMPERATURES:*,
+/* * * AMBIENT:*,23x,FB.2,*C *,F8.2,*F*,
+/* * * PREHEATED COMB. AIR:*,11x,F8.2,*C *,F8.2,*F*,
+/* * * FURNACE FLAME:*,17x,FB.2,*C *,P8.2,*P*,
+/* * * PREDILUTION FLUE GAS:*,10x,F8.2,*C *,F8.2,*F*,
+/* * * DILUTED FLUE GAs:*,14x,FB.2,*C *,F8.2,*F*,
+/* * * DRYING (BIN PLENUM):*,11X,F8.2,*C *,F8.2,*P*,
+/* *,* OUTER FURNACE WALL:* 12x,F8.2,*C *,F8.2,*F*)
FORMAT(*0*,*MASS FLOWS:*,
+/* *,* WIND SPEED, RPB:*,1sx,FB.2,sx,FB.2,* MPE*,
+/* *,* FUEL FEED, KG/H:*,15x,F8.2,5X,F8.2,* LB/H*,
+/* *,* FUEL MOISTURE, WB:* ,12x, F8. 2,
+/- *, * FURNACE AIR SUPPLY, NG/NG-FUEL:* ,FB.2,
+/* * ,29x,*CMM:*,FB. 2, 5x, FB. 2,* CFM*,
+/* *,* EXCESS AIR, PERCENT. * ,11x, F8. 2,
+/* *,* PREDILUTION FLUE GAS, KG/N:*,4x,FB.2,5x,FB.2,* LB/N*,
+/* *,24x,*CMM:*,5x,F8.2,sx,FB.2,* CFM*,
+/* *,* DILUTED FLUE GAS, NG/H:*,Bx,F8.2,sx,FB.2,* LB/H*,
+/* *,20x,*CHM:*,9X,F8.2,5X,F8.2,* CFM*,
+/* *,* DRTING AIR, CMM:*,15x,FB.2,5x,FB.2,* CFM*,
+/* *,* DRTING AIR ABSOLUTE HUMIDITI:*,4x,FB.4)
FORMAT(*0*,*TOTAL HEAT OUTPUT, MJ/N:*,9x,FB.2,sx,FB.2,* MBTU/E*,
+/* *,*T0TAL MEAT LOSS, MJ/B:*,11x,F8.2,5x,Fa.2,* MBTU/N*,
+/* *,*FURNACE EFFICIENCY, PERCENT:*,5X,P8.2,
+/* *,*SISTEM EFFICIENCY, PERCENT:*,6x,FB.2,
+/* *,*ASH ACCUMULATION, KG/H:*,10X,F8.2,5X,P8.2,* LB/H*)
FORMAT(* *,*UNBURNED FUEL ACCUMULATION, NG/N:*,FB.2,5x,
+P8.2,* LB/H*)

FORMAT<1E+,35X,F7.2>

FORMAT<IN+,33X,A10)

F0RMAT<F7.3)

FORMAT(*0*,*TNIS MIxTURE WOULD'NT BURN!*)

F0RMAT(*0*,*C0MBUSTION AIR, KG/H:*)

F0RMAT(* *,*AIR:FUEL RATIO:*)

END

FUNCTION RUTA(QS,RS,CF,TF,R,T>

DIMENSION R(20)

SB-s.669E-8

..‘Q‘Q

 

253

EBG‘SB*T**4

QUOD=(QS-EBG*RS+T/R(18))*R(18)
RUTA=SB*QUOD**4/R(19)+QUOD*(1./CP+1./R(18))
RUTABRUTA-(T/R(18)+SB*TP**4/R(19)*TF/CF
RETURN »

END

c.......................

10

20

FUNCTION SIMPSON(TEM,REACT,GM,GN,C0M,CP)
DIMENSION C0M(20),CP(20),WMOL(20)

DATA (NMOL(R),x=1,8)/44.01,28.01,28.016,46.,30.,18.,32.,2./
MEATso.

D0 20 L=l,7

IF(C0M(L) .EQ. 0.)GOT0 20

N=20

M=(TEM-298.)/N

HALF-H/Z.

NNsN-1

CM=COM(L)*WMOL(L)
SN1=ENTNAL(298.,REACT,GM,GN,C0M,CP,L)
SNZ=ENTHAL<TEM,REACT,GM,GN,COM,CP,L)

Txn=298.

FXB=FXHALF=O.

TxHALF=298.-HALF

D0 10 III,NN

Tan-12m”!

TXHALF=TXHALP+H
FXH=ENTHAL(TXH,REACT,GM,GN,COM,CP,L)+PXH
FXHALF=ENTHAL(TXHALF,REACT,GM,GN,COM,CP,L)+PXHALF
CONTINUE

TxNALFeTXNALF+N
FxHALFsENTHAL(TXNALF,REACT,GM,GN,COM,CF,L)+FXHALF
SIMPSON=<SN1+SN2+2.*FXH+4.*FXHALF)*H/6.
HEAT=SIMPSON*CM+EEAT

CONTINUE

SIMPSON-HEAT

RETURN

END

C*-mamc1'10)l mLEII—smc

C.
c.
c.

DETERMINES PROPERTIES OF AIR AND FLUE GASES PROM
TABLES OF THE PROPERTIES AT DIFFERENT TEMPERATURES

FUNCTION TABLE(VAL,SMALL,DIPF,K,DUMMY)

DIMENSION VAL(20)
DUMsAMINI(AMAx1(DUMMI-SMALL,0.0),FLOAT(N)*DIFF)
I-I.+DUM/DIFF

IF(I.EQ.N+1)I=N .
TABLE-(VAL(I+1)-VAL(I))*(DUM—FLOAT(I-1)*DIFF)/DIFF+vAL(I)
RETURN

END

FUNCTION ENTNAL(T,REACT,GM,GN,COM,CP,L)
COMMON/PROPETT/Ax(30),ACP(30),APR(30),ADEN(30),AV(30),CFMF(10)
COMMON/FLUE/DNC02(20),CNC02(20),DVCO(20),CNCO(20),DVH20(20),

C.

c.
C.

c.

254

+CKHZO(20),DVHZ(20),CKHZ(20),DVN2(20),CKN2(20).DV02(20),CK02(20)

COMMON/SMS/SMLA,SMLFLU,DIPA,DIFLU,KA,KPLU,SMST,DIPST,KST,COMED(IO)

DIMENSION COM(20),CP(20),WMOL(20)

DATA (WMOL(K),K=1,8)/44.01,28.01,28.016,46.,30..18.,32.,2./

THETA=T/100.

SQT=SQRT(THETA)

SQDBTHETA**2.

CP(1)=-.89286+7.2967*SQT-.98074*THETA+5.78352-3‘SQD

CP(2)=16.526-.16841*THETA**.75-47.985/SQT+42.246/THETA**.75

CP(3)=9.3355~122.56/THETA**1.5+256.38/SQD-196.08/THETA**3.

CP(4)=11.005+51.65/SQT-86.916/THETA**.75+55.58/SQD

CP(5)=14.69-.40861*SQT-16.877/SQT+17.899/THETA**1.5

CP(6)=34.19-43.868*THETA**.25+l9.778*SQT-.88407*THETA‘

CP(7)=8.9465+4.8044E-3*THETA**1.5-42.679/THETA**1.5+56.615/SQD

CP(8)=13.505-167.96/THETA**.75+278.44/THETA-134.01/THETA**1.S

ENTHAL'CP(L)/WMOL(L)*4.19002

RETURN

END

SUBROUTINE PROPATY(TAMB)

COMMON/PROPETY/AK(30),AOP(30),APR(30),ADEN(30),AV(30),CFMF(10)

COMMON/FLUE/DVCOZ(20),CKC02(20),DVCO(20),CKCO(20).DVH20(20),
+CKHZO(20).DVH2(20),CKHZ(20),DVN2(20),CKN2(20),DV02(20),CK02(20)

COMMON/SMS/SMLA,SMLFLU,DIFA,DIFLU,KA,KPLU,SMST,DIPST,KST,COMED(10)

DATA (ADEN(K),K81,25)/1.7684,1.413,1.1774,.998,.883,.7833,.705,
+.6423,.5879,.543,.503,.4709,.4405,.4149,.3925,.3716,.3524,.3364,
+.3204,.3076,.2947,.2827,.2707,.2611,.2515/

DATA (ACP(K),K81,25)/l.006,1.005,1.006,1.009,1.014,1.021,1.0295,
+1.039,1.055,1.064,1.075,1.086,1.098,l.11,1.121,1.132,l.l42,1.151,
+1.16,l.17,1.179,1.188,l.197,1.206,1.214/

DATA (AV(K).K=l,25)/7.49,9.49,16.84,20.76,25.9,31.71,37.9,44.34,
+Sl.34,58.51,66.25,73.91,82.29,90.75,99.3,108.2,117.8,128.2,l38.6,
+148.85,159.1,170.6,182.1,193.8,205.5/

DATA (AK(L),L81,25)/18.1,22.27.26.24,30.03,33.65,37.1,4O.38,43.6,
+46.6,49.5,52.3,55.1.57.79,60.28,62.79,65.25,67.52,70.36,73.2,75.7,
+78.2,80.95,83.7,86.4,89.1/

DATA (APR(K),K=1,25)/.74,.722,.708,.7,.689,.683,.68,.68,.68,.682,
+.684,.686,.689,.692,.696,.7,.702,.703,.704,.7055,.707,.706,.705,
+.705,.705/

DATA SMLA,DIPA,KA/200..50.,24/

DATA (DVC02(L),LII,14)/19.6,27.1,33.6,39.3,44.5,49.5,54.,
+58.4,62.5,66.5,70.3,74.,77.6,81.2/

DATA (CKC02(K),K81,14)/25.9,40.7,54.5,67.2,78.9,89.7,99.9,
*109.3,118.3,126.8,134.9,142.8,150.5,158.l/

DATA (DVH20(L),LBI,14)/13.9,21.2,28.5,35.5,42.1,48.4,54.5,60.3,
+65.8,70.8,75.9,80.8,85.4,90./

DATA (CKHZO(J),J-l,14)/37.6,60.7,86.7,114.6,143.6,173.3,
+203.3,232.5,260.8,287.1,313.6,339.3,363.6,387.9/

DATA (DVCO(K),K91,14)/21.9,29.,35.1,40.6,45.6,50.4,
+55.1,59.5,63.7,67.8,71.7,75.5,78.2,82.8/

DATA (CKCO(L),L81,14)/33.2,45.5,57.5,68.9,79.6,89.8,
+99.S,108.6,117.3,125.7,133.7,141.5,149.,156.2/

DATA (DVHZ(M),M81,14)/IO.9,l4.2,17.1,19.8,22.3,24.6,

 

10

15

10

255

+26.9,29.,31..33.,34.9,36.7,38.5,40.3/

DATA (CKHZ(N),N=1,14)/228.,313.,363.,428.,493.,559.,
+625.,689.,752.,814.,874.,932.,989.,1045./ ‘
DATA(DVN2(N),N=1,14)/21.7,28.5,34.4,39.7,44.8,49.5.54.,
+58.3,62.4,66.4,70.2,73.9,77.6,81.l/

DATA (CKN2(N),N=l,l4)/32.9,44.3,55.6,65.5,77.2,87.2,
+96.6,105.6,114.1,122.3,130.2,137.7,145.l,152.2/

DATA (DV02(M),M=1,I4)/25.7,34.1,41.5,48.1,
+54.l,59.7,65.I,70.4,75.4,80.2.84.9,89.4,93.8,98.l/
DATA (CK02(L),L=1,14)/34.9,49.2,62.6,74.8.
+85.9,96.3,106.5,116.5,126.3,135.9,145.3,154.6,163.8,172.8/
DATA SMLFLU,DIPLU,KFLU/400..200.,13./

DATA (CEMP(L),L=1,6)/2840..2500.,2000.,1020.,660.,380./
DATA SMST,DIPST,KST/.S,.5,5/

Y=1.E-6
X31.E-3

DO 10 L=l,l4
AV(L)=AV(L)*Y
DVC02(L)*DVC02(L)*Y
DVCO(L)'DVCO(L)*Y
DVHZO(L)'DVH20(L)*Y
DVHZ(L)'DVHZ(L)*Y
DVN2(L)=DVN2(L)*Y
DVOZ(L)'DV02(L)*Y
AK(L)'AK(L)*X
CKCOZ(L)'CKC02(L)*X
CKCO(L)=CKCO(L)*X
CKH20(L)=CKHZO(L)*X
CKH2(L)'CKHZ(L)*X
CKN2(L)'CKN2(L)*X
CKO2(L)'CKOZ(L)*X

CONTINUE

DO 15 K=15.25
AY(K)‘AV(K)*Y
AK(K)!AK(K)*X

CONTINUE
RETURN
END

FUNCTION PLUPTY(TEMP,N,R)
COMMON/PROPETY/AK(30),ACP(30),APR(30).ADEN(30),AV(30),CPMF(10)
COMMON/PLUE/DVC02(20),CKC02(20),DVCO(20),CKCO(20),DVHZO(20).
+CKHZO(20).DVH2(20),CKH2(20),DVN2(20),CKN2(20),DV02(20),CKOZ(20)
COMMON/SMS/SMLA,SMLPLU,DIPA,DIPLU,KA,KFLU,SMST,DIPST,KST.COMED(10)

IF(K .EQ.
IP(N .EQ
IP(N .EQ
IF(N .EQ.
IP(N .EQ
IP(N .EQ
IP(N .EQ

2)GOTO 10
1)FLUPTY'TAELE(DVC02,SMLPLU,DIFLU,KPLU,TEMP)
2)PLUPTY=TAELE(DVCO,SMLPLU,DIPLU,KPLU,TEMP)
3)FLUPTY'TADLE(DVN2,SMLPLU,DIPLU,KPLU,TEHP)
4)FLUPTY=TABLE(DVHZO,SMLPLU,DIPLU,KFLU,TEMP)
5)PLUPTY¢TABLE(DV02,SMLFLU,DIPLU.KPLU,TEMP)
6)PLUPTY'TABLE(DVHZ,SMLFLU,DIPLU,KELU,TEMP)

c.

256

IP(N .EQ. 1)PLUPTY=TABLE(CKC02,SMLFLU,DIPLU,KFLU,TEMP)

IP(N .EQ. 2)PLUPTY=TAELE(CKCO,SMLPLU,DIFLU,KPLU,TEMP)

IP(N .EQ. 3)FLUPTY=TABLE(CKN2,SMLPLU,DIPLU,KPLU,TEMP)

IP(N .EQ. 4)FLUPTY=TABLE(CKH20,SMLFLU,DIFLU,KFLU,TEMP)

IF(N .EQ. 5)PLUPTY=TABLE(CK02,SMLPLU,DIPLU,KPLU,TEMP)

IP(N .EQ. 6)PLUPTY=TABLE(CKHZ,SMLFLU,DIFLU,KPLU,TEMP)
RETURN

END

SUBROUTINE AIRPTS(TEMA,D,CKS,VISC,PRN,CPS)
COMMON/PROPETY/AK(30).ACP(30),APR(30),ADEN(30),AV(30),CFMF(10)
COMMON/FLUE/DVC02(20),CKC02(20),DVCO(20),CKCO(20),DVH20(20),

+CKHZO(20),DVH2(20),CKHZ(20),DVN2(20),CKN2(20),DV02(20),CK02(20)

COMMON/SMS/SMLA,SMLFLU,DIFA,DIFLU,RA,RFLU,SMST,DIFST,RST,COMED(10)
DcTABLE<ADEN,SMLA,DIFA,RA,TEMA)
CNS=TABLE<AR,SMLA,DIFA,RA,TEMA)
VISC=TABLE(AV,SMLA,DIFA,KA,TEMA)
PRN=TABLE(APR,SMLA,DIFA,KA,TEMA)
CPS=TABLE<ACP,SMLA,DIFA,RA,TEMA)

RETURN

END

FUNCTION HEATP(T,CM,PUBL,XMDB,TAMB)
COMMON/FROPETT/AR(30),ACP(30),APR(30),ADEN(30),AV(30),CFMF(10)
COMMON/SMS/SMLA,SMLFLU,DIFA,DIFLU,NA,NFLU,SMST,DIFST,RST,COMED(10)
XMW=XMDB/(l.+XMDB)

NATERsFUEL*XMw

RNATERsO.

Fw=1.-xMw
FCP=Fw*(.064+.00028*(T-273.15))*17.391+4.174*xMw
HWOODsFCP*(T-29B.15)*FUEL
CRAIRsTABLE<ACP,SMLA,DIFA,RA,T)
HAIR=CM*CPAIR*(T-298.15)

IF(T .GT. 373.)NNATER-NATER*(2000.+2.01*(T-373.))
NEATFsMNOOD+RAIR+BNATER

RETURN

END

FUNCTION TADI(COM,TEST,GN,NM,CMG,CPG,CON,EEv,ENF)
DIMENSION COM<20), CP(20), x<100>

zERO(A.B)-A-B

Bc-ENv-CON

x-z

x<1>-TEST-600.

x(2)s‘1‘ES'r+100.

NMcGM0w(COM,TEST,CPG,CMG,GN)

CONTINUE

R-x+1

EN1-SIMPSON(X(x-2),2. ,NM,GN,COM,CP)
Fx1-zER0(EN1,BC)
EN2-SIMPSON(x(N-1),2.,NM,GN,COM,CP)
Fx2-zER0(EN2,BC)
x(E>-(Fx2*x(N-2)-Fx1*N(N-1))/(Fx2-Fx1)
ERR-ABS(X(K)-X(K-1))/2.

IP(ERR .LE. 1.E-10)GOTO 10

10

10

257

GOTO 5

CONTINUE

TADI=X(K)
WM=GMOW<COM,TADI,CPG,CMG,GN)
ENF=SIMPSON(TADI,2.,WM,GN,COM,CP)
RETURN

END

FUNCTION GMOW(COM,TG,CPG,CMG,GN)
DIMENSION WMOL(20):COM(20),CP(20)
DATA (WMOL(K),K=l,8)/44.01,28.01,28.016,46..30.,18.,32.,2./
DUMMY=ENTHAL(TG,1.,GM,GN,COM,CP,3)
CMG=CPG=O.

DO 10 181,7

IF(COM(I) .EQ. 0.)GOTO 10
CPG‘CP(I)*COM(I)/GN+CPG
CMG=COM(I)*NMOL(I)+CMG

CONTINUE

GMOW=CMG/GN

CPG=CPG/GMOW*4.19

RETURN

END

APPENDIX D

IN-BIN COUNTERFLON DRYING COMPUTER PROGRAM

258

IN-BIN COUNTERFLOW DRYING COMPUTER PROGRAM

PROGRAM COUNTER< INFUT,OUTPUT,DEBUG-OUTPUT)

C***** C O U N T E R F L O W G R A I N D R Y E R M O D E L
C***** F.W.BAKKER-ARKEMA, PROJECT LEADER

C***** E.N. MWAURA, PROGRAMMER

C.....

C***** DESCRIPTION

C***** MAIN PROGRAM FOR SIMULATION OF AN IN-EIN COUNTERFLOW DRYER
c.....

COMMON/MAIN/XMT,THT,RHT,DELT,CFM,XMO,KAB,TOTEN,TOTH20,XMS,
+TOTSP,IPROD,FM,HDF,XMF,TIME,NEQ,NEQl,NOUR,REFHR,REFBU,TOFENY
COMMON/INPT/BPH,GP,INDl,DELX,DEPTH,DBTPR,XTEMPER
COMMON/PRPRTY/SA,CA,CV,CW,RHOP,CP
COMMON/HLATENT/HA,HB,HFG
COMMON/TENT/K1,K2,XMR
COMMON/IFLAGS/J,JM,ICON,DELXM
COMMON /PRESS/PATM,BCFMF,PG
COMMON /NAME/INAME,IPRODU,IPRODUI,IEMC
REAL K1,K2
COMMON/ARRATS/XM(50),RH(50),T(50,2),H(50,2),TH(50,2),GA
COMMON/COUNT/ITERCT,TIN,THIN,HIN
COMMON/COLD/COOLER,SYSTEM
DIMENSION TDRT(50).XMDRT(50)
F(T)=T+459.69
ERH(T,XM) = 1. - EXP(- .38195*(T + 50.)'XM**2)
C***** INPUT CONDITIONS OF DRYER TO BE SIMULATED
HOUR-0.0
TOTEN=O. $TOTNZO'O. STOTSP-O.
REFHR=REFEU30.
TOFENYIO.
COOLER-O.
PRINT 160
READ 200,5YSTEM
PRINT 199,5YSTEM
IF(SYSTEM .GT. O.)GOTO 73
PRINT 74
READ 200,REFHR
PRINT 199,REFER
PRINT 75
READ 200,REFEU
PRINT 199,REFEU
73 CONTINUE

259

1310

260

IF(SYSTEM .EQ. 2.)PRINT 1300
IF(SYSTEM .EQ. 2.)READ 200,COOLER
IF(SYSTEM .EQ. 2.)PRINT 199,C00LER
IP(SYSTEM .LT. 2.)PRINT 4s
IF(SYSTEM .LT. 2.)READ 200,BPC
IF(SYSTEM .LT. 2.)PRINT 199,BPC
NEQ=2

PRODUCT=1.

NEQI=NEQ

IPROD=IFIX<PRODUCT)

CALL DATE<IPROD>

PRINT 201,1NAME,IPRODU,IPRODU1,IEMC
READ 200,TIN

PRINT 199,TIN

PRINT 190 ‘

READ 200,TAMB

PRINT 199,TAMB

PRINT 202

READ 200,RBAMB

PRINT 199,RHAMB
MIsMADBRH<F<TAMB),RBAMB)

PRINT 240,HI

PRINT 230

READ 200,FUEL

PRINT 199,FUEL
IF(TIN.EQ.TAMB.OR.FUEL.EQ.4.) GO TO 5
CALL ABSN(EIN,TAMB,TIN,HI,FUEL,IPROD)
PRINT 250,NIN

GO TO 8

CONTINUE

PRINT 77

READ 200,BI

MINsMI

PRINT 250,NIN

CONTINUE

PRINT 204

READ 200,TEIN

PRINT 199,TRIN

PRINT 205

READ 200,xM0w

PRINT 199,xM0w
NAME1=10HTESTNEIGNT

NAME2=7M, LB/BU
PRINT*,NAME1,NAME2

READ*,TW

PRINT 199,Tw
RHOP=Tw/1.24*(1.-xMOW/100.)
EMO=xM0w/(100.-XMOW)
RHCN=.9*ERH(THIN,XMO)
NINIT=NADBRN<F<TMIN>.RNCN)
CONTINUE

PRINT 207

261

READ 200,XMF
PRINT 199,XMF
XMF=XMF/100.
IF(COOLER .EQ. 1. .AND. HOUR .NE. 0.)GOTO 1307
PRINT 208
READ 200,DEPTH
PRINT 199,DEPTH
1307 CONTINUE
IF(COOLER .NE. 1.)PRINT 203
IF(COOLER .NE. 1.)READ 200,8CFMF
IF(COOLER .NE. 1.)PRINT 199,8CFMF
IF(COOLER .EQ. 1.)BCFMF=30.
IF(COOLER .EQ. 1. .AND. DEPTH .GT. 8.)BCFMF=20.
CALL SECANT(DEPTH,BCFMF,CFM,PG,COOLER)
C***** PRINT HEADER PAGE OF CONDITIONS AND PROPERTIES
GA'60.*CFM/VSDBHA(F(TIN),HIN)
PRINT 215,CFM,GA,XMO,PG
IF(COOLER .EQ. 0.)PRINT 65
IF(COOLER .EQ. O.)READ 200, ANSWER
IF(COOLER .EQ. 0.)PRINT 199,ANSWER
IF(COOLER .EQ. 1.)ANSWER=1.
IF(ANSWER .LE. 0.)GOTO 1307
PRINT 218
READ 200,TIME
PRINT 199,TIME
PRINT 209
READ 200,DBTPR
PRINT 199,DBTPR
HRPT=10.
IF(SYSTEM .EQ. 2.)PRINT 49
IF(SYSTEM .EQ. 2.)READ 200,HRPT
IF(SYSTEM .EQ. 2.)PRINT 199,HRPT
DELPR=DBTPR '
IF(COOLER .NE. 1.)PRINT 212
IF(COOLER .NE. 1.)READ 200,HDF
IF(COOLER .EQ. 0.)PRINT 199,EDF
IF(COOLER .EQ. 1.)HDF=.75
C***** COMPUTE STEP AND ARRAY SIZES
DELX=.25
IF(SYSTEM .LT. 2.)DELXIBPC*1.24/508.938
IND=INT(DEPTH/DELX)
INDIBIND+1
INDII-IND1+1
C***** COMPUTE INLET RH AND INITIALIZE ALL ARRA! POSITIONS NECESSARY
RHIN=RHDBHA(F(TIN).HIN)
RHT=RHIN
DO 1 I31,IND11
XM(I)‘XMO
H(I,1)'HINIT
T(I,1)=TIN
TE(I,1)'THIN
RH(I)'RHCN

c...

c...

1400

25

160
74
75
45
65

1300
77
49
170
190
198
199
200
201

202

262

1 CONTINUE
8(101)‘HI
RH(1)-RHIN
** CONVERT AIRFLOW TO LB/HR
IF(GA .GE. 500.)NC=.363*GA**.59
IF(GA .LT. 500)HC-.69*GA**.49
CON1=2.*GA*CA
CON2=2.*GA*CV
CON3=HC*SA*DELX
CON4=RHOP*CP
CON5=RHOP*CW
DELT=2.*DELX*(CON4+CONS*XM(1))/(CON1+CON2*H(IND1,1))*.9
IF(COOLER .EQ. 1.)DELT=1.
IF(COOLER .EQ. 1.)EOUR-0.
** CALL IBCFLW FOR SIMULATING IN-BIN COUNTERFLOW DRTING
CONTINUE
CALL IBCFLW4(TAMB,HRPT,TDAVE,XMDAVE)
IF(COOLER .EQ. 1.)GOT0 25
PRINT 1300
READ 200,COOLER
PRINT 199,C00LER
IF(COOLER .EQ. 0.)GOTO 25
IP(SYSTEM .EQ. 2.)GOT0 1400
RHCN=.9*ERH(TDAVE.XMD)
NINIT=EADBRE<F<TDAVE>,RNCN)
TIN=TAMB
TEIN=TDAVE
MIsEINaMADBRE<F<TAMB>,RHAMB)
SISTEM=2.
GOTO 1310
PRINT 260
DEP1=0.
FORMAT<*0*,*COUNTERFLOW? N0=2; TES=1., REFILL-0.*)
F0RMAT(*0*,*TIME BETWEEN REFILLS, BOURs:*)
FORMAT(* *,*BUSHELS PER REFILL:*)
F0RMAT(*0*,*BUSHELS PER CYCLB:*)
FORMAT(*O*,*SATISPIED WITH DEPTH/PRESSURE COMBINATION?*,
+/* *,*FEs-1.; NO=O.*)
F0RMAT(*0*,*COOL THE GRAIN? IE5-1., N0-0.*)
F0RMAT(*0*,5x,*ABSOLUTE HUMIDITY,DECIMAL:*)
FORMAT(*0*,5x,*OUTPUT INTERVAL: HOUR;*)
F0RMAT(1x*EQTN FOR LOW TEMP;(1 RUG,2 MISRA,3 SABBAH):*)
F0RMAT(5x*AMBIENT AIR TEMP, F :*)
F0RMAT(I1)
F0RMAT(1N+,40x,F10.4)
FORMAT(F10.2>
FORMAT(*0COUNTERFL0W GRAIN DRIER SIMULATION*/
+* USING THE *A10,A10* EQUATION FOR *A10 /
+* AND EMC BY *A10// .
+* INPUT CONDITIONS :*/
+sx*DRTING AIR TEMP, F :*)
FORMAT(SX*AMBIENT REL HUM, DEC :*

203
204
205
207
208
209
212
213
215

218
230

240
250
260
300

c..

263

F0RMAT(*OFINES FACTOR; 1—2(1 CLEAN, 2 DIRTY):*)
F0RMAT(sx*INLET GRAIN TEMP, F:*)
F0RMAT(5x*INITIAL MOISTURE, W.B.PERC.:*)
.FORMAT(sx*FINAL MOISTURE,W.B.PERC.:*
F0RMAT(5x*BED DEPTE,FT:*)

F0RMAT(5x*OUTPUT INTERVAL,FT:*)
F0RMAT(5x*NYBRID DRYING FACTOR,DEC.:*
F0RMAT(///1x*FAN Is DOWN FOR, :*)

F0RMAT(//* PRELIMINARY CALCULATED VALUEs*//
+* AIRFLOW, CFM/SQ FT *FB.4/

+* DRY AIRFLOW RATE, LB/MR-FTz *FB.4/

+* INLET MC(DRY BASIS DECIMAL) *FB.4,/

+* PLENUM PRESSURE, IN-HZO: *,F8.4)
FORMAT<5x*MAx.DRYING TIME, HR:*)
F0RMAT(sx*TYPE 0F FUEL USED (1=NO.2 FUEL*/
+5X,*2=NAT.GAS; 3-L.P.GAS; 4=BIOMASS):*)
FORMAT(SX,*CALCULATED AMBIENT ABS NUM=*9x,F10.4)
F0RMAT(5x*CALCULATED INLET ABS NUM=*11x,F10.4)
F0RMAT(5x,*TNIS IS THE END OF COUNTERFLOW*)
FORMAT(5X*DRYING AIR TEMP, F :*)

END

c * I...B-C.888888=388338. “m8

 

SUBROUTINE IDCFLW4 (TAMB, HRPT, TDAVE , XMDAVE)

C * .I'muiulu'38-‘88-“888'38888338888388I...“
c . . . . .

DIMENSION RHN(50),CYCLE(50),TDRY(40),xMDRY(40),X(50)
COMMON/MAIN/XMT,THT,RHT,DELT,CFM,XMO,KAB,TOTEN,TOTHZO,XMS,
+TOTSP,IPROD,FM,NDF,xMF,TIME,NEQ,NE01,NOUR,REFHR,REFBU,TOFENY

COMMON/INPT/BPH,GP,IND1,DELX,DBPTH,DBTPR,XTEMPER

COMMON/PRPRTY/SA,CA,CV,CW,RNOP,CP
COMMON/ARRAYS/MGO).RH(50),'1‘(50,2),H(50,2),TH(SO,2),GA
COMMON/HLATENT/HA,NB,NFG
COMMON/IFLAGS/K,JM,ICON,DELXM
COMMON/PREss/PATM,BCFMF,PG
COMMON/COUNT/ITERCT,TIN,THIN,HIN
COMMON/HRS/HR

COMMON/COLD/COOLER,SYSTEM

EXTERNAL ZEROIN

EXTERNAL SOLVE4

DATA PATM/14.696/

DATA RHC,AREA/.999999,254.469/

F(T)-T+459.69

ERHCI'JIM) II 1. - EXP(- .38195*('r + 50.)*XM**2)

C.....

c.....

IF(COOLER .EQ. 1. )REFHR’O .
IF(REFHR .GT . 0 .)SYSTEM=1 .
I'IRPTS'BO .

EMBIN=XDD

DRYDUT‘O .

PWATER‘PBU'O .

264

GLEVEL1=DEPTH
HZOOUTao.
YL=0.0

8380.
CTIME=25./60.
KKIO

CYCLE1=0.
KBU=O

RAB=0

REALT=0.

RC0N=0

DRM=0.
DEEP=REFIN-0.
ITERCT=0
NODESsINT(.3/DELX)+1
RAB=0
IND11=IND1+1
IND2=IND1
TOTBTUW=0.0
TINN=TIN=T(1,1)
RNINeRHDBNA<F<TIN>,HIN)
DPCN=2.*DELx
xMBTstO

c.....

C***** BEGIN TIME LOOP
C.....
40 YL'YL+DELT
IF(CYCLEl .GT. 0. .OR. SYSTEM .GE. 2.)REALT=.1
IF(COOLER .EQ. 1.)REALT=0.
HR‘HR+DELT+REALT
HOUR=HOUR+REALT+DELT
CYCLE2=HOUR
HRPTS'HRPTS+DELT+REALT
REFINBREFIN+DELT+REALT

C.....

C***** COMPUTE MC FOR DEPTHaO
cut...
DEP=DEP1=0.
TN(1,1)-(14.*TE(1,1)+T(1,1))/15.
IF(SYSTEM .EQ. 2.)TN(1,1)-(1.5*T(1,1)+TN(1,1))/2.s
THT=T8(1,1)
XMT=XM(1)
RHT=RHIN
CALL LAYEQ
XM(1)-xMT
H<1.2)-H(1.1)
T(1,2)-TINN
TU(1.2)-TN(1,1)
RNR<1>=RNT*100.

c.....

C***** BEGIN DEPTH LOOP

C.....

265

382

XMVBO.
102 CONTINUE

K33

JM‘J-l

TH<JFZ)=TH(JM,2)

HFG‘HEATLAT(XM(J).TH(J.1))
C.....
C***** USE PREVIOUS VALUE OF H AS INITIAL GUESS
C*****

HJ23H(JM02)

IF(COOLER .EQ. 0.)GOT0 1300

IF(J .EQ. 2)GOTO 1300

THDEL=1.25

THCHKsTH(JM.2)'THDEL

IF(THCHR .LE. T(1.1))H(JM.2)=HJZ=H(1.1)
RH(JM)=RHDBHA(F(T(JM.2))oH(JM.2))

cttttt
C***** CALL SOLVE‘ To COMPUTE TRIAL T.XM.M.RH
c*****
1300 CONTINUE

DIFF‘SOLVE4(HJ2)

C***** CHECK CONDENSATION FLAG
IF(ICON) 31.31.30
30 KCONIKCON+1
GO TO 43
C***** SET LIMITS ON H
31 IF(DIFF) 36.43.32
C***** SOLVING FOR ABSORPTION CONDITIONS
32 HLOW=.5*ERH(THT.XMT)
C***** CHECK FOR FEASIBLE HLOW
33 IF(XMT-EMC(HLOW.THT)) 34.35.35
C*”*** DECREASE BLOW UNTIL XM .GT. EMC
34 HLOW=.5*HLOW
GO TO 33
35 HLOW'HADBRH(F(THT).HLOW)
HBI=H(J72)
KAB‘KAB+1
GO TO 42
C***** SOLVING FOR DRYING CONDITIONS
36 MLOW=H(J.2)
C'**** CHECK FOR SUPERSATURATED GRAIN
IF(XMT-EMC(RHC.THT)) 38.37.37
37 HHIBHADBRH(F(TUT).RMC)
GO TO 42
38 HHII.5*(1.+ERH(TUT.XMT))
C***** CHECK FOR FEASIBLE HUI
39 IF(EMC(HHI.THT)-XMT) 63.41.41
C***** INCREASE UNI UNTIL EMC .GT. NM
63 HBI-.5*(1.+HHI)
GO TO 39
41 HHIIHADBRH(F(TET).NHI)

266

c.....

C***** INITIATE SEARCH FOR H, T, XM
C***** '
42 CALL ZEROIN(HLOW.HHI..00001.SOLVE4)
43 xM(J)=xMT
C***** END DEPTH LOOP
J-J+1
IF(J.LE.IND11) GO TO 102
D0 105 LM=2,IND1
T(LM,1)=T(LM,2)
TH<LM.1)=TH(LM.2)
”(ml 1)'H(LM,2)
m=xuv+m<w>
RHN(LM)=RH(LM)*100.
105 CONTINUE
ITERCT=ITERCT+1
c*****
C***** CHECK IF LONG ENOUGH OR DRY ENOUGH OR TIME TO SAVE VALUES
C***** FOR PRINTING. IF NONE OF THESE GO TO THE BEGINNING OF THE LOOP
Cttttt
IF(RBU .EQ. O)XMBOT=(XM(1)+XM(2))/2.
IF(KBU .EQ. 0.)GTEMP=(TH(3,1)+TH(2,1)+TH<1,1))/3.
XMBTM=XMBOT/(1.+XMBOT)
xMBWBstBTM*100.
C...
CFMTOT=CFM*AREA
xMV=xMV/FL0AT(IND1-1)
m-XMV/(l. +xMV)
IF(SYSTEM .EQ. 2. .AND. HRPTS .GE. HRPT>GOTO 400
IF(HOUR .GE. TIME)GOT0 400
IF(xMV. LE. xMF) GO TO 400
IF(SYSTEM .EQ. 2.)GOTO 40
IF(KBU .GT. 0)GOTO 400
IF(xMBTM .LE. XMF)GOTO 400
GOTO 40
c.
400 CONTINUE
C* SECTION SHIFTS ARRAYS BY DEPTH MODES REMOVED
Ct
IF(SYSTEM .GE. 2.)GOTO 12
IF(KBU .GT. O)GOTO 5
IP(XMBTM .GT. EMF)GOT0 12
s IF(SYSTEM .EQ. 1.)CALL VARYD<HBU,DRMN,DRM,CYCLE,CYCLE1.CYCLE2,RR,
+C00LER)
IF(CYCLE(KK) .GE. .45)GOT0 89
HOUR-CYCLE2-HOUR+CTIME
REFIN=REFIN+CTIME
HR-HR+CTIME
CYCLE(KK)-CYCLE(KK)+CTIME
89 CONTINUE
IF(SYSTEM.EQ.1.)DRYBUT-(GLEVEL1-DEPTH)*AREA/I.24+PBU
DRYBu-DELx*2.*AREA/1.24

3

12
CH:

C.
c.
200

220

403

267

IF(SYSTEM.EQ.0.)DRYBUT=DRYBUT+DRYBU

BPHl=DRYBUT/HOUR

BPHZ‘DRYBU/CYCLE(KK)

IF(XMV .LE. XMF)GOTO 3

IF(KBU .GT. 0)GOT0 40

PRINT 403.KR.CYCLE(KK).DEPTB.PG.CFMTOT.DRYBUT.XMBWB.GTEMP.BPHl.
+BPHZ

CONTINUE

CALL CRSPR(XMAVE.TAVE.THAVE.XMEINW.XMEOUTW.RHK)
POWER=PG*CFM*1.5758E-4/.5
ENERGY8GA*(CA+CV*HIN)*(TIN-TAMB)*HR
FENERGY=POWER*2546.136*HR
HZOOUT=RHOP*2.*DELX*(XMBT-XMBOT)
IF(SYSTEM.EQ.0.)WRMD8RHOP'(XMBIN-XMBOT)*(GLEVELl-DEEP)
IF(SYSTEM.EQ.0.)H200UT=H200UT+WRMD*KK
IF(SYSTEM .EQ. 2.)RZOOUT=0.
WATER=(XMBIN-XMAVE)*RHOP*DEPTH

IF(WATER .LT. 0.)WATER=0.
XMS=XMAVE/(l.+XMAVE)*100.

TOFENY=FENERGY+TOFENY

TOTEN=ENERGY+TOTEN
TOTHZOBHZOOUT+WATER+TOTH20
TOTBTUW=TOTEN/TOTHZO
SECO‘(TOTEN+TOFENY)/TOTH20

HR=0.

IF(SYSTEM .LT. 2.)TDRY(KK)=GTEMP
IF(SYSTEM .LT. 2.)XMDRY(KK)=XMBOT
VDABVSDBEA(F(T(1.1))pH(1.1))
GA'60.*CFM/VDA

PRINT 200. PG.POWER.TAVE

PRINT 220.THAVE.XMS.XMEINW.XMEOUTW.TOTHZO.TOTBTUW.SECO.HOUR

HRPTSso.

FORMAT(//* STATIC PRESSURE,IN H20: *P12.4/
+* HORSE POWER,HP/FT2: *F12.4/

+* AVER.AIR TEMP.,F: *F12.4)
FORMAT(* AVER. PROD. TEMP.,F: *F12.4/
+* AVER. MOISTURE,W.B. PERC.: *F12.4/

+* INLET MOIST. EQUIL.,W.B. PERC.: *F12.4/

+* OUTLET MOIST. EQUIL.,W.B. PERC.: *F12.4/

+* WATER REMOVED,LB/FTz *P12.4/

+* SECO (DRYING ONLY), BTU/LB-H20:*,10x,F12.4/

+* SECO (WITH FAN), BTU/LB-H20:*,13x,F12.4/

+* TOTAL DRYING TIME,HR.: *P12.4)

FORMAT(*0*.*CYCLES*.27X.I3./* CYCLE TIME. HRS*.ZOX.F4.2.

+/

* GRAIN DEPTH. FT*.19X.F5.2./* STATIC PRESSURE. IN-820*

+.118.F5.2.

+/* AIRFLOW. CFM*.19X.F8.2./* GRAIN DRIED. BUSHELS*.IOX.F9.2.
+/1GX.* MCWB*.14X.F5.2./GX.* TEMPERATURE. F*.10X.F9.2.

+/ * AVERAGE DRYING RATE. BU/HR*.4X.F9.2.

268

+/ * DRYING RATE, THIS CYCLE, BU/HR*.3X.F6.2)
IF(REFHR .EQ. 0.)GOT0 91
IF(REFIN .GE. REFHR)CALL REFILL(REFIN,GLEVEL1,xMAVE,COOLER)
IF(REFIN .EQ. 0.)PBU=DRYBUT
91 CONTINUE
XMBIN=XMAVE
IF(DEPTH .GE. DPCH)XMBT=(XM(1)+HM(2)>/2.
C *
C* TERMINATING CONDITIONS
C *
xMFs=xM5/100.
IF(XMFS .LE. XMF)GOTO 73
IF(DEPTH .LE. .995)GOTO 73
IF(DEPTH .GE. 12. .AND. COOLER .NE. 1.)PRINT 193
193 FORMAT(*0*.*BIN 15 FULL, STOP PILLING!*)
IF(DEPTH .GE. 12. .AND. COOLER .NE. 1.)GOTO 73
IF(HOUR .GE. TIME)GOTO 73
GOTO 40
73 CONTINUE
IF(SYSTEM .GE. 2.)RETURN
IND=INDI-1
SUMT=SUMM=0.
DO 55 J=1.KK
SUMT=SUMT+TDRY<J>
SUMM=SUMM+HMDRY<J>
55 CONTINUE
TDAVE=SUMT/xx
XMDAVE=SUMM/KK
DEPTH=2.*DELx*xH
IF(DEPTH .GT. 12.)DEPTH=12.
xMWD=xMDAVE/(1.+xMDAVE)*100.
XMO=XMDAVE
PRINT 1370,xMWD,TDAVE
1370 FORMAT(//* FINAL AVERAGE MOISTURE:*,1Bx,F12.4/
+* FINAL GRAIN TEMPERATURE:*.17X.F12.4)
RETURN
END
c . t
c * Ins:smalmzss-unmmssmswusssnaacuse

SUBROUTINE CRSPR(NMAVE.TAVE.THAVE.XMEINW.XMEOUTW.RHK)

C * 8‘883B88II..8ICIMISIWC:=HW
c. . . . .

DIMENSION RHK(50).NMW(50)
COMMON/ARRAYS/XM(50),RH(50).T(50.2).N(50.2).TH(50.2).GA
COMMON/MAIN/DUM(5).XMO.KAB.SKIP(4).IPROD
COMMON/INPT/BPH.GP.INDI.DELX.DEPTH.DBTPR.XTEMPER

C.....
INDBINDl-l
DBPR=DEP=0.
SUM=SUMT*SUMTH=0.0
IF(DBTPR .GT. DEPTH)GOTO 5
PRINT 50

269

D0 10 J=1.IND1

DEPsDEP+DELx

DBPR=DBPR+DELX

XMW(J)=XM(J)/(l.+xM(J))*100. .

IF(DBPR .GE. DBTPR)PRINT 100,DEP,HMW(J),TH(J,1),T(J,1),H(J,1),RHH

+J)

IF(DBPR .GE. DBTPR)DBPR-O
10 CONTINUE
5 CONTINUE

DO 20 1:2.IND1

SUMT=SUMT+T(I,1)

SUMTH=SUMTH+TH(I.1)
20 SUMtSUM+XM(I)

XMAVE=SUM/FLOAT(IND1-1)

TAVE=SUMT/FLOAT(IND1-1)

THAVE=SUMTH/FLOAT(INDl-l)

XMEINsEMC(RH(1),TH(1.1))

xMEOUT=EMC(RH(IN01),TH(INDl,1))

HMEINW=HMEIN/(1.+HMEIN)*100.

XMEOUTW=XMEOUT/(l.+XMEOUT)*100.

RETURN
50 FORMAT(//* DEPTH M c PROD.TEMP AIR TEMP AIR U.R. REL.HUM.*)
100 FORMAT(1X,F5.2,PB.2.F9.2.P9.2,F9.S.P10.2)

 

END
c..
C*.-.88.:sxssssxsssxsaclns:In:ascsaszasxasnssxssllac
SUBROUTINE DATE(IPROD)
C*I un=s=ssnaclcx
C***** SUBROUTINE USED FOR INITIALIZING CONSTANTS FOR PRODUCTS
COMMON/PRPRTY/ SA.CA.CV.CW.RHOP.CP
COMMON/HLATENT/HA.HB
th
COMMON /NAME/INAME.IPRODU.IPRODU1.IEMC
DATA INAME.IPRODU.IPRODU1.IEMC/IONTROETHOM .IOHTHINLAYER .
+108 CORN .108 DEBOER /
cattwt
C***** INITIALIZE CONSTANTS FOR CORN
SA'239.
CA'0.242
CV30.45
CW=1.0
RHOP'38.71
HA-4.349
BBB-28.25
CP'.268
RETURN
END
c..

c * cuss-“ucsmmuumusmn-uc

SUBROUTINE ABSN(NIN.TAMB.TIN.NI.FUEL.IPROD)

C*833:.38:28.38sass:ass-ssssssssszsscsscsss3.338.383
cttata

270

C***** FUEL=1 STANDS FOR No.2 FUEL
C***** FUEL=2 STANDS FOR NATURAL GAS
C***** FUEL=3 STANDS FOR LIQUID PROPANE GAS
C . . . . .
IFUELsIFIx(FUEL)
IF(IFUEL.EQ.1) A=7.0143E—5
IF(FUEL.EQ.2) A=8.17SE-5
IF(IFUEL.EQ.3) A=7.593E—5
IF(IPROD.EQ.2) CP=0.400
IF(IPROD.EQ.1) CP-0.268
HFUEL=A*(1.+HI)*CP*(TIN-TAMB)
HIN=HI+HFUEL
FUEL=FLOAT<IFUEL)
RETURN
.END
c . .
c * SSSSCCBCRRSIBS“8:323:88888888‘RSSRSCBIBSBB‘BBB

SUBROUTINE SECANT(BINL.BCFMF.CFM.PG.COOLER)

C* :1 sacagsaaa==

c..

 

COMMON PF
DIMENSION x(50)
FACTaBCFMF
x<1)=3500.
x(2)=12999.99
st
4 CONTINUE
H-x+1
Fx1-PCFM(X(K-2).BINL,FACT,C00LER)
Fx2=PCFM(x(H-1),BINL,FACT,C00LER)
X(K)=(Fx2*X(K-2)-PX1*X(K-1))/(FX2-Fx1)
ERROR=ABS(X(K-1)-X(K))/2.
IF(ERROR .LE. 1.E-5)GOTO 10
GOTO 4
10 CONTINUE
CFM=X(K)/254.469‘
PG-PF
RETURN
END

C...

c * .88:I8..RMCBS”B'IMB“C“IB suammm

FUNCTION PCFM(CFMX.BINL.FACT.COOLER)
C*susccxsaasanuecIs-38.23.38xxscsangnzazcn-ssanalucz

COMMON PF

A8PACT'1.0038E-8

886.1304E-4

CIALOG(1.+B*CFMX)

F=l./1.1664

6813000.-CPMX

PG¥A*CFMX**2*BINL/C

PF=(G/601.448)**F ,

IF(COOLER .EQ. 1.)PF’9.39-1.1*ALOG(CFMX)

271

 

 

 

 

 

PCFMBPG-PF
RETURN
END
cat
c* I =8======
FUNCTION HEATLAT(XMC,TH)
C*============I a: 3883::
ct****
C***** FUNCTION USED FOR COMPUTING THE LATENT HEAT OF VAPORIZATION
C*****OF WATER IN THE GRAIN.
ct****
COMMON/HLATENT/HA,HB
IF(XMC .GE. .18)GOTO 10
HEATLAT=(1094.-0.57*TH)*(1.+HA*EXP(HB*XMC))
RETURN
10 HEATLAT=1000.
RETURN
END
C**
C* 38:3:acausaszs-xs====ssx=========sss
SUDROUTINE VARYD(KBU,DRMN,DRM,CYCLE,CYCLE1,CYCLE2,KK,COOLER)
c*88¢cxl as: Bagsssxnscsg
C**

DIMENSION CYCLE(50)
COMMON/MAIN/XMT,THT,RHT,DELT,CFM,XHO,KAB,TOTEN,TOTHZO,XMS,
+TOTSP,IPROD,FM,HDF,XMF,TIME,NEQ,NEQl,HOUR,REFHR,REFBU,TOFENY
' COMMON/INPT/BPH,GP,IND1,DELX,DEPTH,DBTPR,XTEMPER
COMMON/PRPRTY/SA,CA,CV,CW,RHOP,CP
COMMON/ARRAYS/XM(50),RH(50),T(SO,2),H(50,2),TH(50,2),GA
COMMON/HLATENT/HA,HB,HFG
COMMON/IFLAGS/K,JM,ICON,DELXM
COMMON/PRESS/PATM,BCFMF,PG
DATA PATH/14.696/
F(X)=X+459.69
CTIME=25./60.
DELX1=DELX*2.
INDZBINDl
IF(KBU .GT. 0)GOTO 10
KK'KK+1
CYCLE(KK)'CYCLE2-CYCLEl
CYCLElSCYCLEZ
10 CONTINUE
DEPTH=DEPTH-DELX1
DRM=DRM+DELX1
CALL SECANT(DEPTH,BCFMF,CFM,PG,COOLER)
IND=INT(DEPTH/DELX)
IND1=IND+1
IDRMBINDZ-INDI
IF(IDRM .LE. 0)GOTO 35
INDXBIND1+1
DO 15 NC=1,INDX
LLBRC+IDRM

15
35

c**

272

LX3KC+1
T(LX.1)-T(Lx.2)=T(LL.1)
H(LX.1)=H(LX,2)=H(LL,1)
RH(LX)=RH(LL)
TH(KC,1)=TH(KC,2)=TH(LL,l)
XM<KC)=XM(LL)

CONTINUE

KBU=KBU+1

DRMN=DRM

DCONT=2.*DELX

IF(DRM .GE. DCONT)KBU'0
IF(DRM .GE. DCONT)DRM30.
RETURN

END

 

C*=:=I l

FUNCTION SOLVE4(HJZ)

c*suaasxnsnsxsssssssxcsasssxssslu
crawtw

COMMON/ARRAYS/XM(SO).RH(50).T(50,2),H(SO,2),TH(SO,2),GA
COMMON/MAIN/XMT,THT,RHT,DELT,CFM,XMO,KAB,TOTEN,TOTHZO,XMS,

. +TOTSP,IPROD,FM,HDF,XMF,TIME,NEQ,NEQl,HOUR,REFHR,REFBU,TOFENY

COMMON/RRPRTI/SA,CA,Cv,Cw,RHOR,CP
COMMON/INPT/EPN,GP,INDI,DELx,DEPTU,DETPR,xTEMPER
COMMON/IFLAGS/J,JM,ICON,DELXM
COMMON/NLATENT/HA,NE,NEG
COMMON/pRESS/RATM,DCEME,PC
COMMON/ICHECK/INTEGER
COMMON/COLD/COOLER,SYSTEM

DATA pATM/I4.696/

DATA REC/0.9999999999/
R(T)=T+459.69

ICON=O

XTEMEER=O.

IND-INDI-I

c*****

C***** EVALUATE THE CONSTANTS CONl....CON6 AND SET THE INITIAL
C***** GUESS FOR H AND COMPUTE TIME INCREHENT, DELT

C*

IF(GA .LT. SOO.)CHTC=.69*GA**.49
IF(SYSTEM .LT. 2.)DELTA=.5

IF(GA .GE. 500.)CHTC=.363*GA**.59
DELTA-l.

CON1=2.*GA*CA

CON2=2.*GA*CV

CON3=CHTC*SA*DELX

CON4=RHOP*CP

CONSIRHOP*CW
DELT=2.‘DELX*(CON4+CONS*XM(2))/(CON1+CON2*H(IND1,1))*.9
IF(COOLER .EQ. 1.)DELT'1.

ctftfifl

JX‘J+1

273

JMM=JM-l
RRC=0
H(J,2)'HJZ
25 IF(J.EQ.2) GO TO 75
' IE(RKC.EQ.1) GO TO 50
DELH=(H(J,2)-H(JMM,2))/2.
GO TO 100
50 DELR=H(J,2)-N(JM,2)
GO TO 100
75 DELH=O.
IP(KKC.EQ.1) GO TO 50
100 CCON1=GA*(CA+CV*H(J,2))
CCON2=DELx*CNTC*SA
CCON3=<CCON2-GA*CV*DELH)*DELT
CCON4=DELT*NPG*GA*DELU
CCONSsDELx*RROP*(CP+Cw*XM(J))
CCON6=GA*DELT/(RNOP*DELx)
TRETA-(TH(JM,1)+TH(J,2))/2.
IF(COOLER .EQ. I.)THETA=(TH(JM,1)*4.+TN(J,2))/5.
T(J,2)-(CCONI*T(JM,2)+CCON2*TRETA)/(CCONI+CCON2)
IF(COOLER .EQ. I.)T(J,2)=(CCON1*T(JM,2)+CCON2*TUETA)/(CC0NI+CCON2)
THETAT=(T(JM,1)*1.+T(J,2))/2.
TH(J,2)=(CCON3*T(J,2)-CCON4+CCON5*TH(J,1))/(CCON5+CCON3)
IF(COOLER .EQ. 1.)TU(J,2)=(CCON3*TEETAT-CCON4+CCONS*TH(JM,2))/
+(CCON5+CCON3)
IR(TN(J,2) .LT. TB(1,1) .AND. COOLER .EQ. 1.)TH(J,2)=TN(I,I)
IP(T(J,2) .GT. TN(J,2) .AND. COOLER .EQ. 1.)T(J,2)-TH(J,2)
IF(COOLER .EQ. I.)GOTO 3
IF(J .GT. 2)GOTO 3
IE(TR(1,I) .GE. TU(2,2))GOT0 3
IF(SYSTEM .EQ. 2.)GOT0 3
THTR‘TH(111)
TE(1.1)-TH(2,2)
TU(2,2)-TRTR
3 CONTINUE

c.....

C***** COMPUTE RH AND CHECK POR CONDENSATION

C.....
R8(J)-RHDBHA(P(T(J.2)).H(J.2)>
IP(RH(J)-RHC) 150,200,200

c.....

C***** CONDENSATION SIMULATOR

c.....

c.....

200 CONTINUE
TS¢T(J:2)
HS=HADBRH(F(T(J,2)),RHC)
H(J,2)-HS
DHDTBHS-HADBRH(F(T(J,2))-l.,RHC)
AIGA'CV
BR(GA*CA+CCON2)-DHDT*GA*CV*(T(JM,2)-TS)-HS*GA*CV
C=(GA*CA+CCON2)*(DHDT*TS-HS)-DHDT*(GA'CA*T(JM,2)+CCON2*THETA)

274

E(J,2)=(-E+(E*E—4.*A*C)**0.5)/(2.*A)
CCONIsGA*(CA+Cv*N(J,2)>

DELH=H(J,2)-H(JM,2)
CCON3=<CCON2-GA*CV*DELH)*DELT
CCON4=DELT*HFG*GA*DELH
T(J,2)-(CCON1*T(JM,2)+CCON2*TUETA)/(CCONI+CCON2)
TH(J,2)=(CC0N3*T(J,2)-CCON4+CCON5*TH(J,1))/<CCON5+CC0N3)
THTsTH(J,2)

RR<J>=RHDEHA(E(T(J,2)).H<J,2))

IC0N=1

DELXM=XMT-XM(J)

c.....

C***** FIND XM ACCORDING TO THE THIN-LAYER DRYING EQUATION.
C.....
150 XMTBXM(J)
IF(RH(J) .GE. RHC)RH(J)=RHC
THT=TH(J,2)
RHT=RH<J>
CALL LAYEQ
c.....
SOLVE4=XMT-XM(J)+CCON6*(HJZ-H(JM,2))
C* M EQUATION
XMT=XM(J)-CCON6*(H(J,2)-H(JM,2))
H(J,2)=(XM(J)-XMT)/CCON6+H(JM,2)
IF(COOLER .EQ. l.)RETURN
IF(H(2,2) .GE. H(1,1))RETURN
IF(SYSTEM .EQ. 2.)RETURN

 

 

 

 

HTR=H(202)
3(111)‘HTR
END
C*sag 3:1 Bussasnlass-uuug=
c..
C*.:Inssusnscss-l Icacgsnagsssaaxssuanuca:
SUBROUTINE LAYEQ
C* asses=8asaansssscssaasxnesssszssszssssss
C***** DESCRIPTION
C***** SUBROUTINE TO FIND THE MOISTURE CONTENT BASED ON EQUA-

C***** TIONS BY J.M. TROEGER AND P.M. DEL GIUDICE

c.....

C***** USAGE

C***** USED IN THE FIXED BED AND CROSSFLOW MODELS WITH GRAIN
C***** TEMPERATURES BETWEEN 80 F AND 160 F
C.....

COMMON/MAIN/XMC,TH,RH,DELT,CFH,XMD,KA3,TOTEN,TOTHZO,XHS,
+TOTSP,IPROD,FM,HDF,XMP,TIME,NEQ,NEQI,HOUR,REFHR,REPBU,TOFENY
COMMON /NAME/INAME,IPRODU,IPRODU1,IEMC
COMMON/COLD/COOLER,SYSTEM

DATA INAME,IRRODU,IPRODUI,IEMC/IOHTROETUOM ,IOHTHINLAYER ,
+IOU CORN ,ION DEBOER /

C.....

C***** S T A T E M E N T F U N C T I O N S

275

C.....

P1(XM,R,T)=EXP(-2.45+6.42*XM**1.25-3.lS*R+9.62*XM*SQRT(R)+.O3*T-.0
102*CFM)

P2(R,T)'EXP(2.82+7.49*(R+.01)**.67-.Ol79*T)

P3(P,Q)=-.12*(XMD-XME)**(Q+1.)*P*Q

Ql(XM,R,T)=-3.98+2.87*XM-(.019/(R+.015))+.016*T

Q2(R)'-EXP(.81-3.11*R)

TF(P,Q,XO,XF,TO)=P*(XF-XME)**Q-P*(X0 -XME)**Q+TO

M(P,Q,XO,TI,TO)=( (TI-T0)/P+(XO-M)**Q)**(1-/Q)+m

c.....

C***** P R O G R A M
c.....
IF(RH .LT. 0.)RH=.OOOl
C***** CALL READYTH FOR PRELIMINARY CHECKS AND CALCULATIONS
CALL READYTH(TXMO,DELM,XME,IOOPS,XMR)
C***** CHECK ABSORPTION FLAG...IF SET GO TO ABSORPTION SIMULATION
IF(IOOPS-1)1,6,1
1 IF(COOLER .EQ. 1.) GOTO 100
C***** COMPUTE TRANSITION M,P1,Ql, AND FIRST TRANSITION TIME
X1M=.4*DELM+KME
XZM=.12*DELM+XME
TINC=DELT*60.
P=P1(TXMD,RH,TH)
M1(m0.RH.m)
TXBTF(P,Q,TKMO,X1M,O.O)
C***** CHECK IF PRESENT M IS IN FIRST REGION...IF IS IS COMPUTE
C***** EQUIVALENT TIME AND ADD TINC
IF(KMC.LT.X1M) GO TO 3
TI=TF(P,Q,TXMO,XMC,0.0)+TINC
C***** CHECK IF EQUIVALENT TIME+TINC IS LESS THAN TRANSITION TIME..
C***** IF IT IS COMPUTE NEW M AND RETURN
IF(TI.GT.TX) GO TO 2
KMC= HDF*XMN(P,Q,TRMD,TI,0.0)
RETURN
C***** EQUIVALENT TIME+TINC IS IN SECOND REGION--COMPUTE P2, QZ AND
C***** NEW M THEN RETURN
2 P=P2(RH,TH)
Q'QZ(RH)
XMC‘ HDF*XMN(P,Q,X1M,TI,TX)
RETURN
C*'*** M IS NOT IN FIRST REGION--COMPUTE P2, QZ AND SECOND
C***** TRANSITION TIME
3 P=P2(RH,TH)
Q'QZ(RH)
TXl=TX
TX'TF(P,Q,XIM,X2M,TXI)
C***** CHECK IF PRESENT M IS IN SECOND REGION...IF IT IS COMPUTE
C***** EQUIVALENT TIME AND ADD TINC
IF(XMC.LT.X2M) GO TO 5
TI'TF(P,Q,XIM,XMC,TK1)+TINC
C***** CHECK IF EQUIVALENT TIME+TINC IS LESS THAN TRANSITION TIME..
C****' IF IT IS COMPUTE M AND RETURN

276

IF(TI.GT.TX) GO TO 4
KMC= HDF'XMN(P,Q,X1M,TI,TX1)
RETURN
C***** EQUIVALENT TIME+TINC IS IN THIRD REGION--COMPUTE P3, Q3 AND
C***** NEW M THEN RETURN
4 P=P3(P,Q)
Qt-l.
XMC= HDF'XMN(P,Q,X2M,TI,TX)
RETURN
C***** M IS NOT IN SECOND REGION-~COMPUTE P3, Q3, EQUIVALENT TIME+
C***** TINC AND NEW M THEN RETURN
S P-P3(P,Q)
QI-l.
TISTF(P,Q,X2M,XMC,TX)+TINC
XMCg HDF*XMN(P,Q,XZM,TI,TX)
RETURN
ct****
C***** ABSORPTION SIMULATION
C***** FIND NEW M AND INCREMENT COUNTER
6 DIV=-.625'PSDB(TH+459.69)**(.466*RH)*RH*RH*RH
XMC‘ HDF*((XMC-XME)*EXP(DIV*DELT)+XME)
KAB=KAB+1
RETURN
100 IF(IOOPS-l)10,20,10
10 ALMR=ALOG(KMR)
A8-1.86178+0.0048843*TH
88427.364*EXP(-0.03301*TH)
C**** FIND EQUIVALENT TIME BASED ON CURRENT TEMP AND MC
C**** ADD DELT AND SOLVE FOR NEW MC
TI!ALMR*(A+B*ALMR)+DELT
ALMR=(-A-SORT(A*A+4 . 0*B'TI ) )/( 2 . 0*E)
XMC= HDF*(DELM*EXP(ALMR)+XME)
RETURN
20 KABBKAB+1
RETURN
END
C**
C*nnun.ssssznnxnt Inuussacsnnznsscsscnsus:
SUBROUTINE READYTH(TXMO,DELM,XME,IOOPS,XMR)
C*s:xnuznsaasal IE .........
ctuatt

 

 

 

C.....

C***** DESCRIPTION

C***** SUBROUTINE MAKES PRELIMINARY CHECKS AND CALCULATIONS

C***** FOR THIN LAYER EQUATIONS

c.....

c..... USAGE

C***** US!“ WITH LATEQ IN FIXED AND CROSS FLOW DRIER MODELS

c.....
COMMON/MAIN/XMC,TH,RH,DELT,CFM,XMD,KAB,TOTEN,TOTHZO.EMS,
+TOTSP,IPROD,FM,HDF,XMF,TIME,NEQ,NEQI,HOUR,REFHR,REFBU,TOFENY
COMMON /NAME/INAME,IPRODU,IPRODUI,IEMC

C.....
c.....

1

C.....
C.....

2
3

4

C.....

5

C..
c.

277

IOOPS 3 0
COMPUTE EQUILIBRIUM MOISTURE CONTENT, COMPARE TO PRESENT
MOISTURE CONTENT... IF GREATER SET IOOPS = 1
XME 3 EMC(RHITH)
IF(XME - XMC) 2,1,1
IOOPS 3 1
COMPARE PRESENT MOISTURE CONTENT TO INITIAL MOISTURE CONTENT
SET TXMO 3 THE LARGER VALUE
IF(XMO - XMC) 3:494
TXMO 3 EMC
GO TO 5
TXMO 3 HMO
COMPUTE MOISTURE RATIO
DELM 3 TXMO - XME
xMR . (EMC - :IMEVDELM
RETURN
END

 

c.
c.....
C.....
c.....
c.....
C.....
C.....
c.....

c.....

c.....
c.....
c.....
C.....
234
c.....
C.....

300

c.....
c.....

301

302

303 A 3

FUNCTION EMC(RH,T)

E “888888“:

S.F.DEBOER, PROGRAMMER IN ENGLISH UNITS
AND
EDISON W. RUGUMAYO, CONVERTER TO SI UNITS
DESCRIPTION
FUNCTION COMPUTES EQUILIBRIUM MOISTURE CONTENT OF CORN
FROM A RELATIVE HUMIDITY AND TEMPERATURE
CHECK TEMPERATURE TO DETERMINE EQUATION TO BE USED

TC.(T-320 )/108

IF(TC - 112.778) 234,235,235
DEBOER EQUATIONS

CHECK IF RH IS GREATER THAN
IF(RH - .50) 300,300,309
PART ONE RH .LE. .50 ONLY
COMPUTE CONSTANTS
F1 3 -.00070596*TC + .0874496
F2 3 -.00078354*TC + .1188704
F3 3-.00096462*TC + .1474512
51 3 13.838*(-9.*Fl + 6.*F2 - F3)
52 3 13.838*(4.*F3 - 9.*F2 + 6.*F1)
B 3 RH - .17
FIND INTERVAL IN WHICH RH LIES AND COMPUTE EQUILIBRIUM MOISTURE
CONTENT
IF (B) 301,301,302
EMC ' (Sl*RH*RH*RH/l.02 + (Fl/.17 - 51*.02833)*RH)
RETURN
IF(RH - .34) 303,303,304
.34 - RH
EMC 3 (51*A*A*A/1.02 + 52*B*B*B/1.02 + (F2/.17 - 52*.02833)*B +

.50 ..IF SO GO TO SECOND PART

278

+(Fl/.17 - 51*.02833)*A)
RETURN
304 A 3 .51 - RH

EMC 3 52*A*A*A/1.02 + (F3/.17)*(RH - .34) + (F2/.17 - 52*,02333)
+*A
RETURN

C***** PART TWO---RH.GT. .50 ONLY

C***** COMPUTE CONSTANTS

309 F0 3 -.000967l4*TC + .1452064
F1 3 -.00127350*TC + .1848600
F2 3 -.00134082*TC + .2293632
F3 3 -.00192780*TC + .3588280

51 - 13.838*(4.*F0
52 c 13.838*(4.*F3
E - RB - .66
IP(B) 305,305,306
C***** FIND INTERVAL IN WHICH RE LIES AND COMPUTE EQUILIBRIUM MOISTURE
C***** CONTENT
305 A . RH - .49
EMC c Sl*A*A*A/1.02 + (Fl/.17 - 51*.02833)*A + (F0/.17)*(.66 - RH)
RETURN
306 IF(RH - .83) 307,307,308
307 A - .83 - RH
EMC . Sl*A*A*A/l.02 + 52*B*B*B/1.02 + (32/.17 - 52*.02833)*B +
+(Pl/.17 - 51*.028333)*A
RETURN
308 A - 1.0 - RH
EMC - 52*A*A*A/1.02 + (F3/.17)*(RH - .83) + (F2/.l7 - 52*.028333)
1*A
RETURN
cttttt
C***** THOMPSON EQUATION
C***** TC.GE. 113 DEGREES CELCUS
C***** COMPUTE EQUILIBRIUM MOISTURE CONTENT
235 EMC - .01*SQRT((-ALOG(1. - RH))/(.0000382*(1.8*TC + 82.)))
RETURN
END
SUBROUTINE REFILL(REFIN,GLEVELI,XMAVE,COOLER)
COMMON/MAIN/NMT,TUT,RHT,DELT,CEM,NMO,RAE,TOTEN,TOTH20,NMS,
+TOTSP,IPROD,FM,HDF,XMF,TIME,NEQ,NEQI,HOUR,REFHR,REFBU,TOFENY
COMMON/INPT/BPH,GP,INDl,DELX,DEPTH,DBTPR,XTEMPER
COMMON/PRPRTY/SA,CA,CV,CW,RHOP,CP
COMMON/ARRAxS/XM(50),RE(50),T<50,2),E(50,2),TU(50,2),GA
COMMON/HLATENT/HA,HB,HFG
COMMON/IELAGS/R,JM,ICON,DELxM
COMMON/PRESS/PATM,BCFMF,PG
COMMON/COUNT/ITERCT,TIN,THIN,BIN
DATA pATM/14.696/
DATA RHC,AREA/.999999,254.469/
F(T)3T+459.69
REPFT3REFBU*I.24/AREA
DEPTHF3REFFT+DEPTH

9.*F1 + 6.*F2 - F3)
9.*F2 + 6.*F1 - F0)

279

INDF=INT(DEPTHF/DELX)+1

INS=IND1+1

INDEX3INDF+1

DO 10 I3INS,INDEX

XM(I)3XMO

TH(I,1)3TH(I,2)3THIN

T(I,1)3T(I,2)3T(IND1,1)

H(I,1)3H(I,2)3H(IND1,1)

RH(I)=RH(IND1)
10 CONTINUE

SUM30.

IND1=INDF

DO 30 I31,IND1

SUM=SUM+XM(I)
30 CONTINUE

XMAVE=SUM/IND1

GLEVEL1=DEPTH3DEPTHF

PRINT 20,REFBU,DEPTH,HOUR

CALL SECANT(DEPTH,BCFMF,CFM,PG,COOLER)

GA360.*CFM/VSDBHA(F(TIN),HIN)

REFIN=0. '
20 FORMAT(*0*,*BIN REFILLED WITH *,F5.1,* BU TO A DEPTH OF *,F4.1,

+* FT*/* *,*AFTER *,F4.1, * HOURS’)

RETURN

END

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280

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