-------
V = BULK AVERAGE VELOCITY
N= PARTICLE CONCENTRATION
Figure 10. Differential matrix segment.
With collection efficiency defined in the form of Equation (1)
a mass or number balance for particles of a single size can be constructed
in the following form:
Particle Flux In = Particle Flux Out + Particles Collected , (28)
NV = (N + dN)V +
(29)
where — , - F = particles collected per unit length of fiber per unit time
— • — = total length of fiber projected perpendicular to flow in the
ira
unit cross section of matrix
V = V (1 - F)= bulk average velocity upstream of matrix
F = fractional packing density (1 - porosity).
Rearranging Equation (29) and integrating over a matrix of length L
gives
4FLn,
NT = exp
- F).
(30)
where P = penetration
N = particle co
N. = particle concentration at matrix inlet.
N = particle concentration at matrix outlet
32
-------
Equation (30) expresses the theoretical fractional penetration on a
number or mass basis for a single particle size. The applicable value
of nf is selected from Equation (25) or (26). Since the extrapolation
from a single fiber to the total matrix inherently assumes that there
are no particle-to-particle or fiber-to-fiber interactions and contains
O
the idealized 2/ir fiber-orientation factor, the numerical factor 4/ir
should be considered as a first approximation for an "effectiveness
factor" to be verified or determined empirically through experiment.
POTENTIAL APPLICATIONS OF HGMS TO FINE PARTICLE EMISSION CONTROL
Probably the most important parameter to be considered in the
evaluation of potential applications of HGMS is the particle suscepti-
bility. Particles with high magnetic susceptibilities will be most
easily and economically collected by the HGMS method. Significant
emissions of such particulates are found in the iron and steel industry
and the ferroalloy industry. Emissions from various processes in these
industries contain high percentages of FeQ, Fe203, MnO, Mn20,, MngO*,
Cr90o, and other highly paramagnetic compounds. Table 1 lists several
C* O
potential applications with typical reported dust compositions. The
composition obviously varies widely among the different processes.
Within a single process category the composition will vary over a more
narrow range with changes in raw feed and operating conditions. However,
the susceptibility of heterogeneous dust particles cannot be predicted
accurately even when a bulk composition is known. The valence state of
ferromagnetic elements is extremely important because very small per-
centages of ferromagnetic species can produce order-of-magnitude increases
in susceptibility. Fortunately, the susceptibility of small samples can
be determined experimentally using a Faraday balance or similar device,
providing a reasonably simple way to screen potential applications.
33
-------
TABLE 1. POTENTIAL APPLICATIONS OF HGMS TO FINE PARTICLE EMISSION CONTROL
Industrial Process References Typical Dust Composition,
mass basis
Sinter Plant - Highly variable
Blast Furnace [71] 35-50% Fe, 12% FeO, 0.5-0.9% Mn
Basic Oxygen Furnace [71,72] 90% Fe203, 1.5% FeO, 4% Mn304
Open Hearth Furnace [71,72] 85-90% Fe203, 1-4% FeO, 0.5% MnO
Electric Arc Furnace [71,72] 20-55% Fe203, 4-10% FeO, 0.5% MnO
Silico-manganese Furnace [71,72] 4-7% FeO, 30-35% MnO
Ferro-manganese Furnace [71,73] 6% FeO, 34% MnO
Ferro-chrome Furnace [71,73] 7-11% FeO, 3% MnO, 29% total Cr
Particle size also varies widely in different types of dusts, and
the susceptibility can vary with size in a single process due to
compositional changes with size. HGMS theory predicts in general a
decrease in collection efficiency with size, but experimentation with
particular dusts will-be required to determine the minimum feasible size
to collect. Gas temperature may also be an important factor. Trendwise,
particle susceptibilities are proportional to the reciprocal of absolute
temperature assuming Curie law behavior [74]. Gas viscosity increases
with increasing temperature. Both of those trends theoretically reduce
HGMS efficiency at higher temperatures, but the effects could be minor.
Again experimentation will be required to determine the technical and
economic practicalities of the process. Theoretical aspects, however,
can play a strong role in designing experiments and equipment and
understanding experimental results.
34
-------
SECTION 5
EXPERIMENTAL WORK
BASIS OF EXPERIMENT DESIGN
The bench-scale experimental test program was planned to evaluate
the preliminary technical and economic feasibility of applying high
gradient magnetic separation to fine particle control. Existing methods
of fine particle collection for air pollution control are generally
characterized by a limited range of capital costs and energy require-
ments. Although the actual values vary greatly with the specific
technology applied, the type of source controlled, and the efficiency of
particulate collection, many common applications require a capital
investment of roughly $2000 to $8000/m3/s of gas flow ($1 to $4/CFM) for
uninstalled primary equipment. Power consumption frequently ranges from
1.6 kW to 3.2 kW/m3/s of gas flow (1 hp to 2 hp/1000 CFM). Total mass
collection efficiencies of greater than 99 percent are typical, and the
current trend is to design for at least 90 percent collection of sub-
micron particulate. If HGMS is to become a viable alternative to
present control methods, satisfactory performance must be demonstrated
with competitive estimates of capital cost and power requirements or
some other significant advantage over present technology must be exhibited.
To focus the HGMS evaluation on practical operating conditions, the
upper bounds of the above rule-of-thumb capital costs and power require-
ments were adopted as tentative performance goals. With appropriate
design information and assumptions these goals were then transformed
into tentative limitations on major operating variables. The trans-
formation is explained in the following paragraphs.
Capital Cost
In late 1975 when this analysis was first conducted, the largest
commercially operating HGMS systems were iron-bound solenoids with a
2.13-m (7-«) diameter bore and a 50-cm (20-inch) matrix length.
35
-------
According to one vendor of commercial-scale equipment, the cost of a
unit of this type with a field rating of 0.2 T would have been approxi-
mately $240,000 including the magnet, power supply, and cooling equip-
ment [75], This estimate indicated that the superficial gas velocity
would have to be at least 8.4 m/s (1650 ft/min) to keep the capital
•5
cost below $8000/m /s ($4/CFM) with a duty cycle approaching 100 percent.
Power Requirements
The major power requirements of HGMS in a particle/gas system would
be generation of the magnetic field and operation of the gas moving
equipment. To produce a background field,of 0.2 T, the large iron-bound
solenoid referenced above would require approximately 20 kW electrical
input [75]. A unit of the same size has been reported to consume 200 kW
of power to produce a field of 1.0 T [76]. These two figures provided
baseline data that were used to calculate magnet power requirements for
the field/velocity combinations shown in Table 2.
TABLE 2. ESTIMATED MAGNET POWER REQUIREMENTS FOR
SEVERAL FIELD/VELOCITY COMBINATIONS
Background Field
T
0.2
0.2
1.0
1.0
Gas Velocity
m/s ft/min
5 984
10 1968
5 984
10 1968
kW
20
20
200
200
Magnet
kW/m3/s
1.12
0.56
11.23
5.61
Power
hp/1000 CFM
0.71
0.35
7.12
3.57
Fan power requirements are directly proportional to the product of gas
flow rate and pressure drop. With a 60-percent efficient fan a pressure
drop of 2.5 kPa (10 inches of water) is equivalent to a power requirement
of 4.17 kW/m3/s (2.63 hp/1000 CFM).
36
-------
The bench-scale experiment design and execution were strongly
influenced by the results of this analysis. However, since the design
of large electromagnets has been primarily for research usage in the
past (with the exception of the several units now in use in the clay
industry), cost reductions and more energy-efficient designs might be
anticipated should HGMS emerge as a widely used technology. Accordingly
the limiting conditions for the major operating variables were initially
specified at the following levels:
Operating Parameter . Limiting Condition
Superficial gas velocity >5 m/s (984 ft/min)
Pressure drop <2.5 kPa (10 inches of water)
Background magnetic field 90% for submicron particles
>99% for particles greater than 1 ym
Basic oxygen furnace dust was selected as the test dust from the
list of likely candidates because it satisfied three important criteria:
(1) it is associated with a major industrial process, (2) it is difficult
to collect by conventional methods because of the high mass percentage
of submicron particles, and (3) it is one of the strongest candidates for
successful HGMS application because of the typically large ferric oxide
content.
SUMMARY OF PHASE I EXPERIMENTS AND RESULTS
Apparatus and Procedures
The experimental apparatus utilized in Phase I is shown schematically
in Figure 11. Basically, the apparatus consisted of a fluidized bed
system to redisperse the dust followed by a section in which the dust
was diluted with clean air. The dirty air stream was then passed
through the magnetic separator and vented. Samples were extracted
upstream and downstream of the magnetic separator and analyzed to
determine the particle size distribution and concentration.
37
-------
ROOM
AIR -
INTAKE
co
00
HEATER
CONTROL
FILTER
FLUIDIZED
BED
FILTER
DUST FEEDER
DRIVE
O
SCR
MOTOR
CONTROL
HGMS
I ENTRAIN WENT
SEPARATOR
U
t
r
RARTICLE
ANALYZER
J*
3-
BLEED—XI*
DILUTER
Q.
-CXJ-
VENT
F1LTER
C.W.
-><3 I^BLEED
INTAKE
FILTER
COMPRESSOR
Figure 11. Experimental apparatus used in preliminary HGMS experiments.
-------
A dust redispersing system was chosen for source simulation over
artificial particle generators because it allowed the use of an actual
industrial dust at a comparatively high feed rate. The fluidized bed
system was similar in principle, although somewhat different in appli-
cation, to equipment described by Guichard [77], Blann and Moreno [78],
and Willeke and co-workers [79]. The fluidized medium was nickel metal
powder sieved to a size range of approximately 75 to 100 ym. Fluidizing
air was compressed and cooled, passed through an entrainment separator
and submicron filter, and metered into the bed to provide a superficial
velocity of approximately 15 cm/s.
The test dust was obtained from the hoppers of an electrostatic
precipitator which controls emissions from a basic oxygen furnace at a
Pennsylvania steel mill. The dust was fed into the fluidized bed by a
small screw feeder. The fluidized nickel spheres partially deagglomerated
the bulk dust, allowing it to be entrained in the superficial air flow.
The fluidized BOF dust was diluted by mixing with a stream of clean room
air. The incoming air was preheated to 38° C (100° F) to reduce the
relative humidity and filtered to remove some of the background particulate.
To minimize the complexity of the experimental apparatus, all of the
Phase I tests were limited to 38° C.
The dusty air stream was passed through the high gradient magnetic
separator, which consisted of a 35.6-cm long steel canister with an
inside diameter of 8.9 cm. The canister was positioned in the bore of
an iron-bound solenoid where the background magnetic field could be
adjusted in six increments from 0.094 to 0.965 T. The electromagnet and
its associated power supply constituted a commercially available laboratory
HGMS unit which was leased from Magnetic Engineering Associates of
Cambridge, Massachusetts. The filter material was 430 stainless steel
wool similar to that used in previously reported experiments [49].
Sampling and analysis were accomplished using a custom dilution
system and a commercial optical particle analyzer. Sequential sampling
was conducted upstream and downstream of the HGMS apparatus to determine
the collection efficiency as a function of particle size. Since rather
long sample lines were required, information on the larger particles
(3 vm and above) was unavoidably lost.
39
-------
Dilution of the samples was required to avoid coincidence loss in
the optical analyzer. Specifications for the dilution system were
obtained through the courtesy of Southern Research Institute (SoRI) in
Birmingham, Alabama. In the SoRI-designed diluter the sample stream is
fed into a conical vessel where it is mixed in highly turbulent flow
with a filtered stream of dilution air. A dilution ratio of 50:1 was
used throughout these tests. Calibrations by SoRI [80] have shown that
this type of diluter yields representative samples very close to the
predicted dilution ratio for submicron particles. Some transmission
losses experienced by SoRI with 2.3 urn particles during calibration
caused greater uncertainty in the predicted dilution ratio.
Single point, isokinetic sampling was initially attempted in this
preliminary work, but a problem developed because larger particles were
drawn through the sample lines and deposited in the diluter. These
particles would then occasionally become broken up and entrained,
causing spurious counts. To eliminate the problem, it was necessary to
turn the sample nozzles downstream to prevent large particles from
entering the sampling apparatus. This extreme deviation from isokinetic
sampling probably led to some bias in the sampling of even the fine
particles. However, since the upstream and downstream particle concen-
trations were ratioed to calculate a collection efficiency, it was assumed
that the sample bias was eliminated by the ratio (i.e., that the bias
for any given particle size was a constant factor independent of particle
concentration and equal at the upstream and downstream sample points).
The particle analyzer was a Climet Model 208A which uses near-
forward scattering of visible light to count and size particles in
discrete increments. The optical particle analyzer was chosen primarily
because it allowed the measurements to be made in near-real time. The
particular unit used in this study was a special-order model including
an extra increment in the submicron range. Data were extracted for four
optically determined ranges of particle diameter: 0.3 to 0.5 ym, 0.5
to 0.7 ym, 0.7 to 1.0 ym and 1.0 to 3.0 ym.
During each HGMS run three or more particle counts were taken up-
stream and downstream in sequence. A collection efficiency was cal-
culated from each paired data set, and the results were averaged to
40
-------
obtain the efficiency for the particular run. Although the individual
data sets sometimes yielded penetration results differing by as much as a
factor of two, they showed no trend with time; that is, all of the data
were collected before significant loading of the matrix occurred.
Between runs the filter matrix was forward-flushed (aided by manual
rapping) with the magnet off. Occasional duplication of runs showed
that this method of cleaning returned the matrix to its original
condition within the limits of experimental error. One long-term run
was made to observe the effects of uninterrupted filter loading.
Salient Experimental Results
The ranges of important operating parameters are summarized in
Table 3. Figures 12, 13, and 14 show results of several runs that
encompass the ranges of these variables. The dramatic improvement in
collection efficiency upon application of a low magnetic field is most
easily seen in those groups in which zero-field runs were made. In-
creases in the background field further reduced the particle penetration
as expected. Fields higher than 0.308 T were never needed because in
all tests the collection efficiency approached or exceeded detectable
limits with the 0.308 T field.
The effect of gas velocity was less obvious. Efficiencies ex-
ceeding 90 percent were achieved with both low and high velocities.
Increases in velocity seemed beneficial in some test series and detri-
mental in others. At this point in the program the effect of gas
velocity on collection efficiency was not well understood.
TABLE 3. RANGES OF OPERATING PARAMETERS - PHASE I EXPERIMENTS
Magnetic Field, T
Matrix Length, cm
Packing Density
Gas Velocity, m/s
Pressure Drop, kPa
(in. H20)
0-0.308
20.3
0.0088
5.9-9.3
0.7-1.3
(2.8-5.3)
0-0.214
20.3
0.0132
6.1-10.6
1.2-2.9
(4.7-11.5)
0-0.214
20.3
0.0174
5.9-10.2
1.4-3.2
(5.4-12.8)
41
-------
1.0
0.5
0.2
O.I
0.05
0.02
V=8-4 mfy
V = 9.3 m/s
IU
z
UJ
QOI
0-005
O.OO2
aooi
B0=0
0.094T
I i
0.2 0.4 0.6 0.8 I 2
PARTICLE DIAMETER, fj.m
1.0
0.5
0.2
O.I
0.05
<0.02
LU
Z
OOI
0.005
0.002
0001
0.094T
0.2 0.4 0.6 0.8 I 2
PARTICLE DIAMETER^m
1.0
0.5
0.2
O.I
OO5
0.094T
-0.214T
0.308T
O
|0.02
UJ
UJ
a.
OOI
0005
0002
0.001
O.2 0.4 0.6 0.8 I
PARTICLE DIAMETER,
Figure 12. Results of bench-scale HGMS tests with packing density of 0.0088.
-------
V= 6.1 m/s
V=8.4m/s
V=I0.6 m/s
1.0
DC
Ul
z
UJ
a.
0.2
O.I
0.05
0.02
0.01
O.OO5
O.O02
0.00
0.094T
O.2 0.4 0.6 0.8 I 2
PARTICLE DIAMETER,/*m
1.0
05
O2
O.I
005
£ 002
a.
aoi
0005
0.002
O.OOI
B0=0
0.094
0.2I4T
0.2 0.4 0.6 OS I 2
PARTICLE DIAMETER,/im
IX)
0.5
0.2
O.I
0.05
O
0.02
LJ
a.
001
0005
0.002
0.001
OO94T
0.214T
0.2 O4 0.6 0.8 I 2
PARTICLE DIAMETER,p.m
Figure 13. Results of bench-scale HGMS tests with packing density of 0.0132.
-------
1.0
V=5.9 m/s
O.2
0.1
0.05
< Q02
£C
Ul
Ul
a o.oi
0.005
0.002
O.OOI
0.094T
0.2I4T
O2 0.4 O.6 0.8 I 2
PARTICLE DIAMETER, fJ.m
1.0
0.5
0.2
O.I
005
V=8.l m/s
002
UJ
0.01
0.005
0.002
O.OOI
0.094T
0.2I4T
0.2 0.4 0.6 0.8 I
PARTICLE DIAMETER,
1.0
0.5
0.2
0.1
0.05
V=I0.2 m/s
O
2 0.02
UJ
Ul
0.
0.01
0.005
OXX>2
O.OOI
0.094T
i i i
0.2 0.4 O.6 0.8 I
PARTICLE DIAMETER,
Figure 14. Results of bench-scale HGMS tests with packing density of 0.0174.
-------
The important effect of matrix packing density can also be seen by
comparing runs with comparable velocities in Figures 12, 13, and 14.
Increasing the packing density yielded a definite improvement in col-
lection efficiency but also increased the matrix pressure drop.
One long-term run was made to study the variation in sampling
results with time and the transient effects of matrix loading. The
results are shown in Figure 15. The penetration varied without any
pattern for the first eight hours of the test and then began to rise.
After 10 1/2 hours the superficial gas velocity could no longer be
maintained at its original value because of the increased pressure drop,
and the run was stopped after 11 1/2 hours. At the end of the run the
matrix was carefully removed and weighed. The clean weight was 120 g,
and a total of 81.8 g of dust had been collected. Assuming the total
mass collection efficiency was constant over the 11 1/2 hour run (neg-
lecting the small decrease in the last four hours), these results
indicate that the matrix collected approximately 50 percent of its own
weight in dust before the collection efficiency was appreciably affected.
However, during this time the pressure drop across the matrix also
increased by about 50 percent.
Preliminary Economic Estimates
A preliminary economic analysis was prepared based on the results
obtained with the median packing density matrix (F = 0.0132). Satis-
factory collection efficiencies were obtained with the median density
matrix using superficial gas velocities of 6.1 to 10.6 m/s (1200 to 2090
ft/min). The choice of operating velocity involves a trade-off between
operating and capital costs. A high velocity reduces the total flow
area and thus the number of parallel collector modules, but results in a
higher pressure drop and thus greater energy requirements for fan
operation. Lower velocities have the converse effect. Cost estimates
were made for a hypothetical system using either low or high velocity
operation and taking into account the increased fan power requirements
with higher pressure drop. The incremental costs of HGMS modules were
evaluated on the basis of 2.13-m (7-ft) diameter modules each of which
45
-------
CT>
0.08
O.07 -
2.00
- 1.75
-0.25
Figure 15. Bench-scale variation in collection efficiency and pressure drop with matrix loading,
-------
would cost $240,000 (including the magnet, power supply, and heat ex-
changers) and would require 20 kW to produce a 0.2 T background field
[75]. Since information on matrix cleaning was not yet sufficient to
allow a detailed analysis of operating cycles, a 75-percent duty cycle
was assumed (25 percent downtime on each module for matrix cleaning),
which is comparable to commercial clay unit operation [10]. The results
of the preliminary analysis are summarized in Table 4.
Both the cost and power requirement projections shown in Table 4
are slightly higher than the tentative goals stated previously, but
are well within the range of potential feasibility. As mentioned
previously, the cost of large-scale HGMS systems is not yet firmly
established, and the process has not been optimized for gas-stream
application so the cost projections could easily be ±50 percent.
Also calculations indicate that the total energy requirement of the
system could be reduced by using higher magnetic fields with lower
matrix packing density (hence, lower pressure drop). In the low
velocity case presented in Table 4, magnet energization accounts for
TABLE 4. PRELIMINARY ECONOMIC ANALYSIS FOR
HGMS APPLIED TO BOF DUST CONTROL
Superficial
gas Velocity, m/s
(ft/mln)
Estimated Pressure Drop, kPa
Cost of HGMS
Total Power
(in H20)
Modules, $/m /s
($/CFM)
3
Requirements, kW/m /s
(hp/1000 CFM)
Low Velocity
6.1
(1200)
1.2
(4-7)
14,700
(6.94)
2.9
(1-8)
High Velocity
10.2
(2010)
2.5
(9.9)
8,800
(4-15)
4.7
(3.0)
approximately 32 percent of the total power requirement, and in the high
velocity case, approximately 12 percent.
47
-------
At this stage in the program it became apparent that more detailed
experimental results were needed on a larger scale system to expand the
understanding of the process and to provide more extensive and reliable
results on which to base economic projections.
DESCRIPTION OF PHASE II APPARATUS AND EXPERIMENTS
General Guidelines For Scale-Up
The bench-scale experiments provided a baseline of information and
experience from which to design the larger system. The following guide-
lines for the design of both the pilot-scale apparatus and the experi-
ment resulted.
(1) Contractual agreement stipulated a minimum system capacity of
0.235 m /s (500 CFM). To insure that this obligation could be
met even if operating velocities had to be reduced in the
larger system, a 30-cm diameter matrix canister was specified.
(2) A more systematic experiment was designed to allow analysis of
variance and direct comparison of runs that differed by only
one operating parameter.
(3) Operation at four velocities in the nominal range of 5 to 10
m/s was planned in an attempt to more accurately determine the
velocity effect since it has an important impact on system
optimization.
(4) The matrix packing density was reduced to yield lower pressure
drops and a more equitable distribution of projected power
requirements between the magnet and fans. Two matrix densities
(F = 0.005, 0.010) were chosen for evaluation.
(5) A magnetic field capability of 0.5 T was specified. Zero-
field operation was also planned to aid in the comparison of
experimental data and theory.
48
-------
(6) Two different types of dust were specified to broaden the
information on potential applications. BOF dust from the same
source as the bench-scale test was selected as one dust. A
second dust was obtained from the hopper of a roof-system
baghouse that controls emissions from electric arc furnaces at
a Pennsylvania steel mill.
(7) Allowance was made for a matrix length of up to 30 cm. Two
matrix lengths (L = 15, 30 cm) were selected for testing.
(8) Plans were made for limited testing at an elevated temperature
to allow preliminary evaluation of temperature effects.
(9) Selection of operating parameter levels was purposely directed
toward lower efficiency operation than that observed in the
bench-scale tests to allow better observation of parameter
effects. When collection efficiencies exceed 90 percent
(penetration = 0.1) good repeatability of experiments is more
difficult and parametric effects are more likely to be masked
by random error.
(10) Cascade inertia! impactors were specified as the primary means
of particle size determination since their range of particle
size information is broader than the available optical system
and since impactors are more widely used in field sizing
studies. Also the theoretical filtration and HGMS models are
based on aerodynamic behavior of spherical particles, which is
the same basis on which impactors are calibrated and operated.
The optical system was also retained to provide preliminary
results in near-real time and to allow observation of transients
during extended runs.
Apparatus and Procedures
To avoid the duplication of costly equipment that was already
available, the pilot plant was designed to operate as a slipstream off
the existing Particulate Aerodynamic Test Facility located at the EPA
Technical Center in Research Triangle Park, North Carolina. This
facility basically consists of a low-speed, closed-loop wind tunnel.
Incorporated with the wind tunnel is a dust feed and redispersion
49
-------
system that operates on the same fluidized-bed principle as the system
used in the bench-scale experiments. (In fact, the bench-scale system
was designed as a scaled-down model of the EPA facility.) The wind
tunnel is also equipped with a baghouse that results in once-through
operation with respect to dust. The entire system is described in
detail elsewhere [81,82], and the description will be repeated here only
to the extent that it pertains to the Phase II experiments.
The experimental apparatus is shown schematically in Figure 16.
The fluidized bed in this system had two separate stages. The lower-bed
was 20 cm in diameter and contained 2.5 cm of 6-mm glass beads. It was
operated at incipient fluidization to break up the dust initially as it
was fed in by the screw feeder. The dust then sifted up through a
screen to the 40-cm upper bed, which contained approximately 2.5 cm of
no. 9 lead shot and 2.5 to 5 cm of 700-ym glass beads. The larger
diameter of the upper bed lowered the gas velocity so that the glass
beads were fluidized but not elutriated, and the lead shot was near
incipient fluidization. The upper bed further broke up the dust ag-
glomerates. Some early effort was devoted to optimizing the gas flow
rate and bed depths with an objective of maximizing the percentage of
fine particles in the output, but little success was achieved. The bed
depths were finally set at the above values, and the fluidizing gas flow
was set at roughly 0.14 m /s (300 CFM). These conditions provided the
most trouble-free operation and yielded a steady particle output with
•3
roughly 1 percent smaller than 1 pm (p = 4.5 g/cm ).
From the fluidized bed the dust was blown into the 61-cm diameter
wind tunnel where it was mixed with dilution air coming from the bag-
house and wind tunnel fan. A slipstream of the dusty air was then drawn
off the wind tunnel into the pilot-scale HGMS apparatus. The stream
passed through a cyclone separator, which was designed to provide a 50-
percent cut point (D5Q) of approximately 5 ym (p = 4.5 g/cm3). The
combination of superisokinetic flow at the slipstream intake and scalping
by the cyclone reduced the percentage of large agglomerates in the gas
stream. The characteristics of the stream exiting the cyclone varied
50
-------
TO .
BAGHOUSE'
WIND TUNNEL
TEMPERATURE
AND FLOW o
INDICATORS 0--
BOOSTER
FAN
BLAST
GATE
DOWNSTREAM
I TEST PORTS
HGMS
UPSTREAM
TEST PORTS
CYCLONE
SEPARATOR
CLEAN
FROM
BAGHOUSE
FLUIDIZED
BED
DUST
FEEDER
FLUIDIZING
AIR
Figure 16. Schematic representation of pilot-scale HGMS facility.
-------
from run to run with the slipstream flow rate, but typically the mass
o
median diameter was on the order of 2 to 4 ym (p =4.5 g/cm ) and the
" 3
total mass concentration was in the range of 50 to 100 mg/m .
From the cyclone this dusty gas stream flowed up through a baffle
which was designed to break up the vortex. It then flowed past the
upstream sample ports and into the HGMS canister, which was 30 cm in
diameter and 35 cm long in the direction of flow. From the HGMS
canister the clean gas passed by the downstream sample ports; through a
return bend; through a blast gate, which was used for flow control;
through a booster fan; and back into the wind tunnel where it rejoined
the remainder of the dusty gas and continued around the wind tunnel loop
to the baghouse.
The matrix inside the HGMS canister was subdivided into four 7.5-cm
long sections by expanded metal screens* which were connected by three
threaded rods running the length of the canister in the direction of
flow (see Figure 17). A small air-operated vibrator was attached
through the duct wall via a rubber grommet to one of the threaded rods.
To construct a 15-cm long matrix, the first and third sections (count-
ing from the upstream or inlet end) were packed with the appropriate
mass of randomly oriented 430 stainless steel wool of the same type used
in the bench-scale experiments. Imbedded in the first section were two
20-cm diameter loops of 0.6-cm (1/4-inch O.D.) copper tubing. The loops
each had numerous 1.5-mm perforations on their underside and were con-
nected to a source of dry, compressed air via a vertical section of
tubing running up the side of the canister and through the duct wall.
The third section contained one of these loops of perforated tubing. To
construct a 30-cm matrix, all four sections in the canister were employed,
and perforated tubes were added to the second and fourth sections.
After each run the vibrator and compressed air were used in tandem to
shake and blow the collected particulate off the matrix.
To conduct an experimental run the wind tunnel was started up and
the slipstream fan was started. The desired nominal velocity was set
by adjustment of the blast gate from a reading on a stationary pi tot
tube located between the canister and the blast gate. A 16-point pi tot
52
-------
TO EXTERNALLY
MOUNTED
VIBRATOR
CANISTER
TO COMPRESSED
AIR SOURCE
EXPANDED
METAL SCREENS
CONNECTING
ROD
\ PERFORATED
/— TUBES
/
Figure 17. HGMS matrix construction and cleaning apparatus
53
-------
traverse (eight points on each of two perpendicular diameters) was then
run at the downstream test ports. A special test was conducted early in
the program to compare the upstream and downstream average velocities as
well as 16- and 32-point traverses. The results are shown in Table 5.
The apparent larger flow rate on the downstream side suggested the
possibility of in-leakage, but no leakage could be detected. It was
more probably due to a bias on the upstream pi tot due to the slight
swirling flow component remaining from the cyclone. In any event the
results were considered close enough to justify the use of the 16-point
downstream traverse as the routine method of flow determination.
When the pi tot traverse was completed, the dust feed was started,
and the magnet was energized if appropriate. A 10- to 15-minute time
interval was allowed to insure steady-state output by the dust system,
and then sampling was begun.
The cascade inertial impactors utilized were MRI Model 1502 units
manufactured by Meteorology Research, Inc., Altadena, California. Prior
to use the MRI substrates were coated with a thin layer of Apiezon L
gease dissolved in toluene, baked at 120° C for 2 hours, desiccated
overnight or longer, and tared. Because of the relatively small duct
diameter (30 cm) and upward direction of gas flow, extractive sampling
was judged to be preferable to in-stack sampling. The sample nozzles
and probes were specially constructed using recommended techniques to
minimize particle losses [83]. The impactors were supported in a
horizontal position and were moved carefully to avoid dislodging
collected particulate. Eight-point traverses were used on the upstream
side and 12-point traverses on the downstream side, both divided between
TABLE 5. COMPARISON OF PITOT TRAVERSE RESULTS
16-point traverse 32-point traverse
Upstream 7.57 m/s 7.39 m/s
Downstream 7.84 m/s 7.82 m/s
54
-------
two perpendicular diameters. Downstream sampling was started first and
lasted 96 minutes. Upstream sampling was then begun and lasted 32 or 48
minutes. Each of the impactors was operated with a nozzle and sample
rate combination calculated to yield isokinetic sampling at the down-
stream average velocity.
After the runs the impactors were carefully dismantled, and the
substrates were desiccated overnight before weighing. The probes were
washed with acetone, and the probe washings were dried and also desic-
cated overnight. All of the weights were determined to 0.01 mg. A
Mettler Model H51AR balance was used for the runs conducted prior to May
17, 1977, and then a new Perkin-ETmer microbalance, Model AD-2Z was
placed in service for the remainder of the tests. The probe washings
averaged 7.7 percent of the total collected mass on the upstream side
and 13.0 percent on the downstream side. This particulate matter was
most probably made up of larger particles lost in the probe bend. Data
from the first two impactor stages (particle diameters of about 5 ym
and above) were not included in the data analysis.
The cut points (DCQ) of the impactor stages were calculated using
the latest recommended EPA procedures [84], including published calibration
2
factors for the MRI impactor stages [85]. Particle densities of 4.47 g/cm
3
for the BOF dust and 4.61 g/cm for the EAF dust were determined from
helium pyknometer measurements (courtesy of Mr. Ray Grote of EPA) and
o
were used in the D5Q calculations. A value of p = 1.0 g/cm is frequently
used in D5Q calculations and yields what is known as the aerodynamic
particle diameter. However, in this case since the volume susceptibility
was used in magnetic calculations, it was decided that the Stokes diameter
based on the measured particle density would be a more representative
measure of particle size. Hence, measured particle densities and Stokes
diameters were used throughout the calculations.
Differential particle size distributions were calculated to reduce
the upstream and downstream data to the same concentration basis. During
the sampling, identical nozzles were used on both sample trains, and
sample rates were adjusted so that D50's of corresponding upstream and
downstream stages were equal (±7 percent). This technique avoided the
55
-------
use of curve fitting procedures and allowed the calculation of particle
penetrations from a direct ratio of upstream and downstream differential
distributions at corresponding impactor stages.
Optical particle counts in the ranges 0.3 to 0.5 ym, 0.5 to 0.7 ym,
and 0.7 to 1.0 urn were also made using the same dilution system and
Climet instrument employed in the bench-scale tests. In these experi-
ments, however, the sample nozzles were turned upstream, and samples
were drawn at a constant nozzle velocity of approximately 5.0 m/s for
all runs. This resulted in sample rates from approximately isokinetic
in the low-velocity runs to 50 percent isokinetic in the high-velocity
runs. Four-point, single diameter traverses were made on both the
upstream and downstream sides. To prevent large agglomerates from
entering the diluter and causing spurious counts, a small cyclone with a
o
DgQ of approximately 1 ym (p = 4.5 g/cm ) was located at the inlet of
the diluter. The subisokinetic sampling rates, the single diameter
traverse, and the presence of the cyclone could have created systematic
errors in the optical data; hence, the impactor data were utilized for
primary data analysis. The optical system was utilized to ensure..that
no upsets occurred during impactor sampling, to provide approximate
efficiency results in near-real time, and to study transient effects
during two special long-term matrix-loading tests.
When each run was completed, the dust feed and gas flow were shut
off. The magnet was deenergized, and the vibration and blowing system
was used to dislodge the collected particulate from the matrix and to
cause it to fall down into the cyclone hopper. The gas flow was then
restarted and the matrix pressure drop was compared to the value ob-
served before the run to ensure that the matrix was clean. If necessary
the cleaning procedure was completed. Difficulties were sometimes
encountered in returning the matrix pressure drop to its original value.
Frequently a net gain in pressure drop of 25 to 125 Pa (0.1 to 0.5 in
F^O) was observed. The small vibrator and blowing system were evidently
not completely adequate to overcome the combination of residual magnetism
and mechanical entrapment of dust. The matrix loading tests conducted
later indicated that this problem probably was not severe enough to
seriously bias the efficiency results, but the cleaning system should
still be upgraded before any future work is undertaken.
56
-------
Experiment Design
The overall layout of the pilot-scale experimental program is shown
in Table 6. Practical constraints prevented a complete randomization of
the experiments. Changes in matrix length or packing density necessi-
tated a time-consuming dismantling of the system so the different matrix
configurations were separated into blocks, which are separated by the
dashed lines in Table 6. With each matrix, the magnetic field was
always incrementally increased, since returning to a lower field could
have introduced errors because of hysteresis. In each set of three or
four runs with the same matrix and same field, the velocities were
randomly ordered. A few preliminary runs were made to debug the equip-
ment and establish operating procedures. The actual test plan was
conducted between April 19, 1977, and July 18, 1977. In Table 6 a run
number of 4201, for example, indicates the first run on April 20, while
4202 indicates the second run on the same day.
57
-------
TABLE 6. EXPERIMENT DESIGN FOR PILOT-SCALE HGMS
Run No.
4191
4201
4202
4211
4271
4272
4281
4282
5031
5032
5041
5042
5091
5101
5102
5111
5121
5131
5161
5171
5181
5182
5191
5192
5201
5202
5242
5251
5252
5261
5262
5271
5272
6011
6012
6021
6022
6031
6032
*6281
7111
7112
7121
7122
7131
7132
7141
7142
7151
*7181
Parti cul ate
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
BOF
EAF
EAF
EAF
EAF
EAF
EAF
EAF
EAF
EAF
EAF
Packing
Density
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.010
0.010
0.010
0.010
0.010
0.010
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
0.005
Matrix
Length
cm
15
15
15
15
15
15
15
15
15
15
15
15
15
15
15
15
15
15
15
15
15
15
15
15
15
15
30
30
30
30
30
30
30
15
15
15
15
15
15
15
15
15
15
15
15
15
15
15
15
15
Magnetic
Field
T
0
0
0
0
0.05
0.05
0.05
0.05
0.10
0.10
0.10
0.10
0.10
0.10
0.10
0.10
0.20
0.20
0.20
0.20
0
0
0
0.05
0.05
0.05
0
0
0
0.05
0.05
0.05
0.05
0.10
0.10
0.10
0.10
0.10
0.10
0.40
0
0
0
0.10
0.10
0.10
0.20
0.20
0.20
0.40
fias
Velocity
m/s
5.59
8.21
6.71
9.79
5.49
8.06
6.74
9.93
8.47
5.63
6.89
10.20
10.78
8.52
7.02
5.46
8.00
6.71
5.70
10.16
9.74
5.94
7.38
5.71
7.84
7.17
5.16
7.65
9.24
7.53
10.06
10.70
6.56
5.68
8.52
10.30 '
6.84
6.76
7.06
8.24
8.45
7.85
10.96
7.88
11.12
9.16
7.53
9.31
10.76
8.21
Temperature
°C
31.1
30.6
34.4
30.8
25.3
32.5
29.1
33.9
30.9
35.3
31.9
36.7
108.9
109.4
109.4
109.7
32.8
33.0
32.8
33.9
30.0
39.4
32.8
37.8
32.8
38.3
33.3
29.3
36.9
30.9
37.0
31.1
35.9
30.9
37.2
32.5
37.2
30.3
37.0
35.4
34.7
41.4
36.4
40.6
35.8
41.1
35.6
40.8
33.1
36.7
Matrix removed from canister, manually cleaned, and replaced before this run.
58
-------
SECTION 6
PILOT-SCALE EXPERIMENTAL RESULTS
Since they predict the effects of individual parameters on col-
lection efficiency, the theoretical models discussed in Section 4
provide a logical framework within which experimental results can be
analyzed. In this section supportive data on matrix and particle
characteristics are first presented, and then the comparison of indi-
vidual runs with theory is presented graphically. This procedure not
only provides a convenient means of presenting all of the experimental
data but also gives an overview of the extent to which observed experi-
mental results can be explained. The effects of individual operating
parameters are then discussed in more detail.
SUPPORTIVE MATRIX AND PARTICLE DATA
The 430 stainless steel wool used as matrix material in these
experiments is commonly used in HGMS systems. It is a very non uniform
material, as evidenced by the photomicrographs shown in Figures 18, 19,
and 20. Some fibers are reasonably well approximated by the idealized
cylindrical geometry while others are more ribbon-shaped and still
others are shaped nearly like a coil. The theoretical investigation of
Stekly and Minervini [59] implies that to a first approximation the
assumed cylindrical geometry should give reasonable predictions provided
the diameter of the cylinder approximates the mean projected width of
the actual fibers. However, Oberteuffer's [1] conclusions on the
importance of matching fiber size to particle size to achieve maximum
magnetic force implies the possibility of different effective mean fiber
diameters for different particle sizes. An impractically large number
of photomicrographs would be required to define a meaningful distri-
bution of fiber sizes and shapes for incorporation into the theoretical
expressions. Therefore, a single radius of 25 urn was selected for use
59
-------
CTl
O
68X
100 urn
68X
lOOym
Figure 18. Photomicrographs of stainless steel wool matrix fibers,
-------
60X
100
200X
50 urn
Figure 19. Photomicrographs of stainless steel wool matrix fibers,
-------
CT1
I •
I 1
150X 100 ym
150X
100
Figure 20. Photomicrographs of stainless steel wool matrix fibers,
-------
in the models. This figure is reasonably consistent with the photo-
micrographs and results in good agreement between theoretical and
experimental results.
Experimental magnetization measurements were also made on seven
small samples of the steel wool material. The range and average of
these measurements are shown in Figure 21. The variability of the
magnetization results from inhomogeneities in the material, cold-worked
effects, and the fact that the different samples had different shapes.
In a total matrix a much higher applied field is required to saturate
the steel wool. Also shown on Figure 21 are three points on the cal-
culated in-matrix magnetization curve for a matrix packing density on
the order of 0.01 [86]. The matrix,eventually reaches saturation with
an applied field on the order of 5.6 to 6.4 x 105 A/m (7 to 8 kOe) [41,
63].
With the low fields used in this work, the matrix was never satu-
rated. To utilize the theoretical model, the near field correction
parameter was thus calculated as
y0M
A = 2B(31)
o
where M = estimated in-matrix magnetization from Figure 21.
The values of the parameter are given in Table 7 for the four fields
used in the experiments.
TABLE 7. VALUES OF THE NEAR FIELD CORRECTION PARAMETER
Background Magnetic Field, tesla A =
0.05 2.39
0.10 2.26
0.20 1.98
0.40 1.54
63
-------
MEASURED
SMALL-FIBER
MAGNETIZATION
ESTIMATED
IN-MATRIX
MAGNETIZATION
) O.I
1
0.2
i
0.3
A/m X IO'6
i i
0.4
i
0.5
i
0.6
i
0.1
0.2 0.3 0.4
TESLA
APPLIED FIELD
0.5
0.6
0.7
0.8
Figure 21. Magnetization curves of steel wool material and in-situ matrix.
64
-------
Experimental determinations of particle magnetization were also
obtained for each of the two dusts. The results are shown in Figures 22
and 23 for the BOF dust and EAF dust, respectively. For convenience the
results are shown in the mass cgs units in which they were determined.
The particle susceptibility is determined from these data by the equation'
x = 4*Ppa/H0 (32)
where x = volumetric particle susceptibility, dimensionless SI units
Pp = particle density = 4.47 g/cm3 (BOF) or 4.61 g/cm3 (EAF)
a = specific magnetization, emu/g
HQ = applied field, Oe.
Figures 22 and 23 show that the susceptibility of both dusts is rela-
tively large and is not constant with applied field. Rather than being
weakly paramagnetic, both dusts exhibit relatively high permeability.
Note that the magnetization of the BOF dust is about twice that of the
EAF dust at any given field.
COMPARISON OF RESULTS WITH THEORETICAL MODELS
The no-field experimental data were fit to an equation of the form
With the fiber radius a = 25 ym and all of the operating parameters
substituted into Equation (33), two adjustable parameters remain: E,
**
the matrix effectiveness factor, and A, the lumped adhesive force factor
that is required for the calculation of ec- The matrix effectiveness
factor would be affected by variations in the tortuous path that the gas
stream must follow through the matrix. Under the idealized assumption
of randomly oriented fibers which do not interfere with each other
aerodynamically, a theoretical value of 4/ir is predicted. In reality,
the value of E that is required to obtain a good fit to the data is also
influenced by the assumed mean fiber radius. The adhesive force
65
-------
o>
z
o
H- 30
N
h-
Ul
z
o
o
UJ
Q.
1.0 2.0 3.0
APPLIED FIELD.kOe
4.0
Figure 22. Magnetization curve of basic oxygen furnace dust.
66
-------
i.o 2.0 ao
APPLIED FIELD, kOe
Figure 23. Magnetization curve of electric arc furnace dust.
67
-------
factor was shown in Section 4 to have a predicted value in the range of
0.002 to 0.1 N/m.
The no-field experimental data for the two dusts are plotted in
Figure 24 through 27. Theoretical curves are shown for comparison. For
both dusts an effectiveness factor of 0.22 was used in Equation (33).
An adhesive force factor of 0.003 provided a reasonable fit to the EOF
data, and the lower EAF collection efficiencies were fit with an ad-
hesive force factor of 0.0005. The results suggest that the rolling
coefficient of friction is a more realistic approach and indicate that
the EAF dust does not adhere as strongly to the wire fiber surface as
the BOF dust does. This is presumably a result of differences in the
,composition of the two dusts which give rise to different Van der Waals
force characteristics.
Lawson's version of the HGMS trajectory model [64,65] was used to
correlate the results of the experiments conducted with a magnetic field
applied to the matrix. Lawson's results presented in Figure 8 were
calculated with a value of the near field correction, A, equal to 0.88.
Table 7 indicates that a more realistic value of A for the lower field
data is 2, while the 0.4 T data correspond to A = 1.5. To accomplish
the transformation of Lawson's results to the appropriate near field
correction, both Lawson's and Watson's [48,51] results were used. First
it was verified that Lawson's and Watson's results are identical at low
values to the Stokes number by interpolating Watson's results to obtain
the curve for A = 0.88 and comparing the capture radius under identical
conditions, recalling that 2AWK = Vm/VQ. Then using Watson's results at
A = 2.0, selected values of the capture radius, RC, were positioned on
Figure 28 along the line log K = -2. The R contours were then extended
V*
with slopes of -1 out to values of log K where the negligible inertia
assumption is no longer valid. For RC = 1 the contour remains straight
for all K. The contours for RC = 4 and 8 were completed accurately using
additional Lawson results in the form log (AW) vs. log K [65] and his
arguments that the separate parameter A has no effect on large R .
\*
68
-------
en
UJ
2
UJ
CL
o
o:
10
05
0.2
0.1
0.05
0.02
0.01
1.0
0.5
O2
Ol
Q05
O02
RUN NO. 4191
BOF
V« 5.59 m/s
LMScm
F - O.OO5
E = O22. A'=O.O03 H/m
O.OI
RUN NO. 4202
BOF
V=67l m/s
L = l5cm
- F=O.OO5
E;O.22,A=O.OO3 N/ra
_L
01
_L
02 05 102 5
PARTICLE DIAMETER, fitn
ui
1.0
O5
0.2
0.1
"Q05
0.02
0.01
1.0
0.5
0.2
5
0.1
005
Q02
O.OI
10 0
RUN NO. 42OI
BOF
V = 8.2I m/t
L= 15cm
F = O.OO5
E=0.22.A' =
N/m
UJ
z
UJ
"-
RUN NO. 4211
BOF
V=9.79 m/s
F- O.O05
t-- O.O03, A1: 0 003 N/m
I 1
I 02 0.5 10 2 5
PARTICLE DIAMETER, p. m
10
Figure 24. Zero-field pilot-scale results with BOF dust (F=0.005, L=15 cm).
-------
1.0
05
0.2
01
O05
002
0.01
N/i
OJ O2 O5 TO 2
PARTICLE DIAMETER,
1.0
O.5
0.2
O.I
0.05
0.02
OOI
1.0
0.5
0.2
O.I
005
0.02
0.01
10
- F= O.OK)
E=022, A'=OO03N/m
Ql
02 05 1.0 2
PARTICLE DIAMETER,
RUN NO. 5191
BOF
8 =O
V 7.38 ra/$
L=I5 cm
_ F = 0.010
E=O.H2, A'=O.OO3 N/m
01 02 0.5 1.0 2
PARTICLE DIAMETER,
10
10
Figure 25. Zero-field pilot-scale results with BOF dust (F=0.010, L=15 cm).
-------
0.0
01
RUN NO. 5242
BOF
B0=0
V= 5.25mA
LOO cm
= O.OO5
. A'=O.OO3 N/m
E=0.2Z, A'= 0.003 N/m
0.2 0.5 1.0
PARTICLE DIAMETER,
1.0
0.2 0.5 10
PARTICLE DIAMETER,
B0=o
V= 9.24 ro/s
t=3Ocm
05 1.0 2
PARTICLE DIAMETER,
Figure 26. Zero-field pilot-scale results with BOF dust (F=0.005, L=30 cm)
-------
--4
ro
1 -\S
0.5
o 0.2
H
a:
ui 0.1
• •I
UJ
o.
0.05
0.02
0.01
•
-
-
RUN NO. 7111
EAF
B0=0
V= B.45m/>
L=l5cm
F = 0.005
E= O.22, A'=O.O005 N/m
1 1 i * i
01
0.2
0.5
I.O
PARTICLE DIAMETER,
I.Oi
2
o
0.5
02
O.I
ui
z
OO5
002
001
1.0
Q5
02
o
t-
Ql
ui
2
Ul
o.
0.05
0.02
10
QOI
RUN NO. 7112
EAF
V = 7.65ra/»
L-IScm
E = 0.22 ,A=0 0005 N/ra
1
1
O.I 0.2 0.5 1.0 2
PARTICLE DIAMETER,
10
RUN NO. 7i2i
EAF
80=0
V= 10.96m/s
L=l5cm
- F=OOO5
£ = 022, A'=O.OO05 N/m
0.1 0.2 0.5 1.0 2
PARTICLE DIAMETER, /.
10
Figure 27. Zero-field pilot-scale results with EAF dust (F=0.005, L=15 cm)
-------
m
The remaining contours were completed by extrapolating with approxi-
lately the same curvature as Lawson's curves at corresponding values of
RC- The extrapolated sections of the contours are shown as dashed lines
in Figure 28. They have not yet been verified by trajectory calculations
and should be regarded as approximate. A similar transformation was
accomplished to model the two runs which were conducted at 0.4 T using a
value of A = 1.5.
Theoretical curves for all of the HGMS runs were calculated using
the expression
P ~ exp
The theoretical curves and experimental da.ta are presented in Figures 29
through 39. The runs are presented basically in the order in which they
were conducted except that the two 0.4 T runs are shown together in
Figure 39. A fiber radius of 25 ym was used both in the determination
of R from Figure 28 and where it appears explicitly in Equation (34).
\f
Equation (34) has one adjustable parameter, E, the matrix effectiveness
factor. A value of E = 0.09 provided a good fit to the BOF data, while
the EAF data indicated a lower value of E = 0.07.
It is reasonable to expect that the effectiveness factor might not
be the same for the two dusts or for the same dust under with-field and
no-field conditions. As argued above, the value of E is influenced by
the assumed value of mean fiber size, by the validity of the assumed
geometric orientation of the fibers, and by the assumption of no inter-
ference among neighboring fibers. In general all of these factors could
be different with respect to aerodynamic and magnetic phenomena, resulting
in different values of E for the with-field and no-field cases.
The differences in effectiveness factors for the two dusts with the
field energized could be related to the matrix loading effects discussed
by Luborsky and Drummond [53,56] and by Cowen, et al. [52,57,62]. The
results of the matrix loading tests presented in Figures 40 and 41
provide additional information. Both of the matrix loading tests were
(34)
73
-------
-2
A = 2.0
G = 0
i -
-2
-I
0
logK
Figure 28. Lines of constant capture radius with inertia! forces included and A=2.0.
74
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o
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Ul
Q-
1.0
05
0.2
O.I
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RUN NO. 5O3I
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V= 8.47m/4
L= IS cm
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PARTICLE DIAMETER,
Figure 30.
Pilot-scale HGMS results with BOF dust (BQ=0.10T).
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(WRTICLE DIAMETER,
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PARTICLE DIAMETER,
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Figure 31. Pilot-scale HGMS results with BOF dust at elevated temperature.
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00
001
0.1
02 0.5 1.0 2
PARTICLE DIAMETER,
O.I
0.2
0.5
PARTICLE
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DIAMETER
10
Figure 32.
Pilot-scale HGMS results with BOF dust (BQ=0.20T).
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NOTE SCALE CHANGE
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PARTICLE DIAMETER,
0.01
0005
0.002
0.001
10
Figure,33. Pilot-scale HGMS results with BOF dust (F=0.010)
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0.5
0.2
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PARTICLE DIAMETER
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Figure 34. Pilot-scale HGMS results with BOF dust (L=30 cm).
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0.5
0.2
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RUN NO 6021
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PARTICLE DIAMETER,
10 O.I 0.2 0.5 1.0 2
PARTICLE DIAMETER,
10
1.0
0.5
0.2
2 0.1
a
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TO.OOM T
o.oooo
0.1
02 0.5 1.0 2
PARTICLE DIAMETER
10
Figure 35. Pilot-scale HGMS results with BOF dust (duplicates of earlier runs).
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2
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10
O.I 02 0.5 1.0 2
PARTICLE DIAMETER,
10
J-
0.1
0.2 0.5 1.0 2
PARTICLE DIAMETER.
10
Figure 36. Pilot-scale HGMS results with BOF dust (replicate runs),
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fc
tu
1.0
0.5 -
0.2
0.1
00
Go
O.O5
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RUN NO. 7122
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B0=O.IOT
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RUN NO. 7131
EAF
BO=OJOT
V: 11.12 m/s
L= IS cm
F= O.OO5
0.1 02 05 1.0 2
PARTICLE DIAMETER,
1.0
10
O.I
02
05
PARTICLE
J_
_L_
1.0 2
DIAMETER,
10
0.5 -
O.I
02 0.3 1.0 2
PARTICLE DIAMETER,
Figure 37. Pilot-scale HGMS results with EAF dust (B =0.10T),
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10
05
Q2
2 o,
UJ
z 0.05
ui
QL
0.02
00
001
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0.2
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RUN NO. 7142
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Ho=0.20T
V*9.3I m/s
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f = O.OO5
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1.0 2
DIAMETER,
10
0.1
03
05
PARTICLE
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0.5
02
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RUN NO. 7ISI
EAF
B0= 0.20T
V= 10.76 m/s
LMScm
F = O.OO5
1.0 2
DIAMETER,
[ 0.0057
U.
10
0.1 02 05 1.0 2
PARTICLE DIAMETER,
10
Figure 38. Pilot-scale HGMS results with EAF dust (BQ=0.20T),
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B0= 0.4OT
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J_
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PARTICLE DIAMETER,
10
001
- 0.005
- 0.002
0.01
0.001
0.5 1.0 Z
PARTICLE DIAMETER,
10
Figure 39.
Pilot-scale HGMS results with both dusts at BQ=0.40T.
-------
0.10 r-
co
en
0
100
200 300
TIME, MINUTES
400
600
-.2.25
- 2.00
S.
JH
a:
a
- 0.25
600
Figure 40. Pilot-scale variation in BOF collection efficiency and matrix pressure
drop with matrix loading.
-------
00 i
0.8
0.7
0.6
0.5
-------
begun with a clean matrix, and operation was continued for several
hours. The optical particle counter was used to make periodic col-
lection efficiency determinations. With the BOF dust there was no
evidence of deterioration in the collection efficiency as the matrix was
loaded. It has been suggested that this behavior, which was also
observed in the bench-scale tests, might be due to the high suscepti-
bility of the dust [87]. Since the particles have a high suscepti-
bility, they may actually distort the magnetic field sufficiently to
maintain the large field gradients in the interstitial spaces rather
than fill the gradient region. Thus the aerodynamic wire radius in-
creases slightly but the magnetic collection mechanism is not grossly
deteriorated. Eventually the accumulation of particles and the pressure
drop become so large that sloughing of collected particles will probably
occur. In the bench-scale test, which was conducted with a field of
0.094 T, the deterioration was apparently beginning when the matrix had
collected about half its own mass (see Figure 15). In the pilot-scale
test, which was conducted with a field of 0.4 T, the matrix had col-
lected nearly twice its own mass of particulate when the run was stopped,
and no deterioration in collection efficiency had been observed. It is
possible that large particles which were not within the range of the
optical system were being sloughed off at this stage.
The matrix loading test with EAF dust yielded distinctly different
results. Figure 41 indicates a gradual deterioration of collection
efficiency from the start of the test. The susceptibility of the EAF
dust is not as large as that of the BOF dust so the particles apparently
do not contribute as significantly to the production of field gradients.
When the EAF loading test was stopped after 450 minutes, the matrix had
collected about three times its own mass in particulate. (Although the
collection efficiency was lower in the EAF test, the inlet mass concen-
tration was higher; hence, more dust was collected.) Since the col-
lection efficiencies determined by inertia! impactors represent integrated
values over 2-hour runs, the theoretical capture radius calculated for
-------
clean fiber conditions is larger than the actual EAF dust capture radius.
In the theoretical equation the lower matrix effectiveness factor provides
a correction.
It should be noted that nearly all of the data correspond to
capture radii that are in the extrapolated portion of Figure 28.
Considering that the extrapolations are a possible source of error, the
overall fit of the experimental data to the theoretical expression is
quite good. There is a fairly consistent trend to overpredict the
collection efficiency of the larger particles, particularly in the low-
field runs. Assuming that the extrapolated capture radii are not
significantly in error, there are at least two possible explanations for
this problem. If the collected particles do contribute new regions of
high gradient, the gradients are most effective over a dimension
approaching the particle size. Considering Oberteuffer's arguments on
the relative size of collector and particle [1], this effect is not as
strong for larger particles, and they are thus more vulnerable to
reentrainment (or failure to be collected) as the matrix becomes loaded.
Secondly, the possibility of particles bouncing off the wires or knocking
others off has not been included in the model. This phenomenon could
possibly be significant with the larger, high momentum particles.
DISCUSSION OF INDIVIDUAL PARAMETER EFFECTS
Analysis of individual parameter effects requires comparison of two
or more experimental runs. Many combinations are possible because of
the experiment design. In this section representative comparisons are
made, and the results are discussed. Before drawing conclusions from
these comparisons, it is important to consider the repeatability of
experiments conducted under identical conditions. Because of time and
cost limitations it was not possible to duplicate all runs. Instead a
set of replicates was run at one set of conditions in the central range
of the operating parameters. Figure 42 shows the results of these
replicate runs. The solid symbols represent the results of three
consecutive runs with the same matrix. In each case the field strength
was set at 0.1 T. The temperature ranged from 30.3 to 37.2 ° C and
89
-------
o:
I-
LU
z
UJ
Q.
1.0
0.5
0.2
O.I
0.05
0.02
0.01
0.005
0.002
0.001
RUN NOS. 5041,6022,6031,6032
BOF
B0=O.IOT
V= 6.76-7.06 m/s
L= 15cm
F= 0.005
T= 30.3-37.2 °C
O.I 0.2 0.5 1.0 2
PARTICLE DIAMETER,
10
Figure 42. Comparison of collection efficiency results of replicate runs,
90
-------
the velocity from 6.76 to 7.06 m/s. These ranges indicate the limits of
controllability of the two variables. The open symbols are the results
of a test run 4 weeks earlier with a different matrix which was sup-
posedly identical. The field strength was also 0.1 T, the temperature
was 31.9° C, and the velocity was 6.89 m/s.
Differences in the three runs with the same matrix reflect random
error due to variations in dust characteristics, fluctuations in the
dust feed rate and size distribution during individual runs, errors
associated with the sampling and weighing process, and probably other
unknown factors. The additional run with another matrix reflects all of
the above sources of error plus the possibility of matrix-to-matrix
variation. It is apparent that the matrix variation was significant.
Three other sets of duplicates were run across the same two
matrices at three different velocities. Pooled standard deviations for
the replicates in Figure 42 plus the three sets of duplicates are reported
in Table 8. Dividing the standard deviation by the mean penetration in
each case gives the coefficient of variation, a measure of the variability
as a fraction of the mean penetration at each diameter. Statistics were
not calculated for particle sizes less than 0.4 ym since data in this range
could not be obtained during all runs.
TABLE 8. POOLED STATISTICS ON PARTICLE PENETRATION
Nominal Particle Diameter, ym
Standard Deviation of Penetration
Coefficient of Variation
Degrees of Freedom
0.4
0.100352
0.229
5
0.8
0.047474
0.421
6
1.8
0.07766,5
0.994
6
3.6
0.169624
1.055
5
Figure 43 shows the important practical relationship among matrix
pressure drop, matrix length, packing density, and superficial gas
velocity. The data were collected at the beginning of the first several
runs with each matrix and thus represent clean matrix conditions. The
91
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3.0
-s
°M
E
u
X
cc
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u.
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2.0
Ul
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a.
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1.0
0.8
0.6
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01
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t_ 0.4
0.3
0.2
O.I
A
F= 0.010
F= 0.005
3 456 8 10
SUPERFICIAL GAS VELOCITY, m/s
20
30
Figure 43. Correlation of matrix pressure drop with gas velocity,
matrix packing density and matrix length.
92
-------
straight lines running through the data were obtained from linear
regression of all of the data shown expressed .in logarithmic form.
The equation is
U~= 13.07 F1'32 V1'79 (35)
with the pressure drop expressed in cm H20, the matrix length in cm,
and the velocity in m/s. The pressure drop data also provide an ad-
ditional indication of matrix variation among runs. Of particular note
are the filled circles on the lower line, which are the data for the 30-
cm matrix. Assuming that the linear relationship between pressure drop
and matrix length is correct, these data suggest that the 30-cm matrix
was more porous than the others, possibly due to leakage around the
extra cleaning rings or other construction anomalies.
The experimental collection efficiency results support the contention
that the construction of the 30-cm matrix might have been faulty.
Figure 44 compares the penetration results of the 15-cm and 30-cm
matrices at the same velocity for both the no-field and with-field
cases. Theoretical arguments suggest that in either case doubling the
matrix length should result in the penetration being squared. For
example, a penetration of 0.2 would be improved to 0.04. The improve-
ment with the longer matrix is not that dramatic in either case shown in
Figure 44. Since there is a question about the effective porosity of
the 30-cm matrix, the quantitative effect of matrix length remains
uncertain. Increasing the matrix length definitely improves collection
efficiency but perhaps not as much as the theory suggests.
Matrix density effects at one velocity are shown in Figure 45, with
and without a magnetic field. Again theoretical considerations imply
that doubling the matrix density should square the penetration. Analysis
of Figure 45 reveals approximately that quantitative result in the no-
field case. With the magnetic field applied, the improvement in efficiency
achieved by doubling packing density is even greater than predicted. In
other comparable runs the improvement in efficiency was at least equal
to that predicted.
93
-------
IQ
o
h-
cc
z
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0.
1.0
0.5
0.2
O.I
0.05
0.02
0.01
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= l5cm
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SYMBOL RUN NO.
o 4202
A 5251
fe,T V,m/s T,«C
0 6.71 34.4
0 7.65 29.3
0.001
J_
_J_
_L
_L
O.I 0.2 0.5 1.0 2
PARTICLE DIAMETER,
10
1.0
05
0.2
O.I
z 0.05
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tr
UJ0.02
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L=30cm
SYMBOL
O
RUN NO.
4281
5272
B0,T
0.05
0.05
V,m/s
6.74 29.1
6.56 35.9
_L
I
0.2 0.5 1.0 2
PARTICLE DIAMETER,
10
Figure 44. Experimental effect of matrix length on collection efficiency with and without
an applied magnetic field.
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2
IU
1.0
0.5
0.2
O.I
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SYMBOL
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A
RUN NO.
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5191
i
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0.1 0.2 0.5 1.0 2
PARTICLE DIAMETER, fj.m
F= 0.010
V, m/s T,°C
6.71 34.4
7.38 32.8
_J I
10
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0.5
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SYMBOL
O
RUN NO
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5202
0.05
0.05
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0.1 0.2 0.5 1.0 2
PARTICLE DIAMETER,
10
Figure 45. Experimental effect of matrix packing density on collection efficiency with and without
an applied magnetic field.
-------
The effect of increasing magnetic field is shown in Figure 46 for
both dusts. The particular runs were selected because they spanned the
widest range of fields. The effect of increasing field was not as
dramatic as in some of the bench-scale runs because the packing density
was much lower in the pilot tests. The experimental scatter also masks
the effects to some extent. The reversed position of the 0.2 and 0-4 T
EAF data was contradicted by the optical particle analyzer and was
probably due to sampling error since there is no apparent reason for
this behavior.
Despite the scatter the beneficial effects of increasing field are
apparent. The theoretical effect of increasing magnetic field is not
straightforward. To a first approximation the capture radius is pro-
portional to the product of the particle and wire magnetizations, and
Figures 21, 22, and 23 show that all of the magnetizations are dependent
on applied field to some power less than 1. As both the particles and
wires approach saturation, the beneficial effects of increasing field
become less significant. Hence beyond some point dependent on the
particular dust, it would be more beneficial to adjust other operating
parameters in order to improve collection efficiency.
Velocity effects are the most complex of all theoretically. In
both the no-field and with-field cases, velocity has two compensating
effects. Without a field, increasing velocity increases the Stokes
number and thus enhances the probability of a particle colliding with a
wire. Unfortunately, it also increases the probability of the particle
being reentrained. In the practical HGMS system the effects of velocity
are perhaps best analyzed with reference to the contour plot of Figure
47. Velocity appears as a squared term in the denominator of W and a
linear term in the numerator of K. Hence starting from any point on the
log W vs. log K plot, increasing velocity implies that one move down two
arbitrary units and to the right one unit. Normally this will result in
a reduction in the capture radius as illustrated by the movement from
point 1 to point 2 on Figure 47, corresponding to a doubling of velocity.
However, in the extrapolated portion of the plot when the negative slope
96
-------
IXD
0.05 -
O
o:
2
Ui
Q_
BOF DUST
RUN NO. BQ.T
0
0.05
0.10
0.20
0.40
0.02 -
0.01 -
0.005 -
0.002 -
0.001
01
0-2 0-5 1.0 2
PARTICLE DIAMETER,
z
o
<
cr
H-
tu
z
UJ
Q.
005 -
0.02 -
0.01 -
0.005 -
0.002 -
0.001
0.2 0.5 1.0 2
PARTICLE DIAMETER,
10
Figure 46. Experimental effect of applied magnetic field on the collection efficiency of both dusts.
-------
en
o
-I
-2
A=2.0
G = 0
-2
0
log K
Figure 47. Illustration of the theoretical effect of increasing gas velocity.
98
-------
of the low capture radius contours becomes larger in magnitude than 2,
increasing velocity actually results in an increase in capture radius.
Physically this is due to beneficial inertia! effects of particles
approaching in the projected area of the fiber.
Figure 48 illustrates the experimental effects of velocity. These
plots are not typical in the strict sense of the word because a few
results showed contradictory effects. In general, however, increasing
velocity has no strong effect on the smaller particles, which correspond
to the extrapolated portion of Figure 47. Higher velocities tend to be
detrimental to larger particle collection when the capture radius is
greater than 1. The same behavior is true with no field apparently due
to the greater tendency of larger particles to be reentrained.
Three of the four elevated temperature runs indicated the possibility
of a beneficial temperature effect especially with the larger particles.
The fourth comparison at the highest velocity showed no significant
effect. The theoretical arguments predict a very small decrease in
collection efficiency when the temperature is raised from 30° C to 110°
C. The collection matrix was particularly difficult to clean after the
four higher temperature runs, suggesting that for some reason the
particles adhered more strongly to the matrix. Industrial gas cleaning
applications of HGMS would be more economically attractive if the
process could be operated at temperatures well above the 110° C used in
these experiments. If the observed improvement in efficiency can be
shown to be real in more extensive testing, it will be an important
result. However, the observed difficulty in matrix cleaning would have
negative implications.
OPERATING DIFFICULTIES
Operating problems were encountered in three separate areas of the total
pilot-plant system. The first two were associated with the dust generation
and wind tunnel system. While these problems have no direct implication
toward application of the HGMS process, they do represent limitations on
future pilot-plant work that could be significant. The third difficulty
was associated with the HGMS matrix cleaning system and suggests the
need for considerably more developmental work in this area.
99
-------
o
o
z
o
UJ
z
UJ
a.
1.0
0.5
0.2
O.
0.05
0.02
0.01
0.005
0.002
0.001
SYMBOL
O
D
_L
RUN NO.
4191
4202
4201
4211
_L
0
0
0
0
J_
V,m/s
5.59
6.71
8.21
9.79
_L
31.1
34.4
30.6
30.8
O.I O.2 0.5 1.0 2
PARTICLE DIAMETER,
10
z
o
oc
»-
UJ
z
UJ
a.
1.0
0.5
0.2
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0.05
0.02
0-01
0.005
0.002
0.001
SYMBOL
O
a
RUN NO.
4271
4281
4272
4282
B,T
0.05
0.05
0.05
0.05
V, m s T,°C
549 25.3
6.74
8.06
9.93
29.1
32.5
33.9
0.1 0.2 0.5 1.0 2
PARTICLE DIAMETER,
10
Figure 48. Experimental effects of gas velocity on collection efficiency with and without an applied magnetic field.
-------
The capabilities of the fluidized bed dust generation system have
been studied in some detail using fly ash as the test dust [82]. With
the iron oxide dusts used in these tests, however, several new diffi-
culties were encountered. First, dust feed rates were limited to about
2 kg/hr. Higher feed rates were attempted so that higher dust concen-
trations could be studied, but the dust would accumulate on the lower
bed and eventually force shutdown. The percentage of submicron par-
ticles in the system output was also much lower than desired. Several
different combinations of fluidizing media and fluidizing gas flow rate
were tried with little improvement. Workable conditions were finally
achieved by keeping the feed rate as high as possible without causing
plugging, reducing the wind tunnel dilution air flow to as low a level
as possible, and using the slipstream cyclone to remove large agglo-
merates ahead of the H6MS canister. With the present facilities the
system cannot be used to study the collection of higher concentrations
of submicron particles that would be more typical of industrial conditions.
The second problem area involved the higher temperature work. As
the wind tunnel is presently constructed, steam can be added to the air
to raise the humidity, but there is no means of determining how much
steam is being added or what the absolute humidity is when the gas
temperature is greater than 50° C. The humidity can be determined with
a sample train, but it cannot be set at a desired point before the run
is made. In this experimental work the humidity was allowed to fluc-
tuate with the ambient conditions during the lower temperature work, and
small amounts of steam were added in the higher temperature runs so that
the humidity would not be extremely low. Better control of the steam
addition and real-time monitoring of the air humidity will be needed
before the system can be used to simulate industrial gas streams at an
elevated temperature and a realistic percentage of water vapor.
' Since the bench-scale HGMS matrices were cleaned by manual rapping,
the quantitative requirements of matrix cleaning were not well defined
when the pilot unit was designed and constructed. As a result, the
vibrator and air blowing system installed proved inadequate. Transmission
of vibrational energy from the externally mounted vibrator to the matrix
101
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was poor, and the air blowing system was not sufficient to completely
overcome the residual magnetism holding the particles to the matrix
wires. These same concepts can probably be applied successfully however,
with the experience gained from the pilot-plant operation and the
quantitative background provided by the reentrainment modeling studies.
The complex relationship involving matrix length, packing density,
particle susceptibility, particle size, and particle mass concentration
needs considerably more attention. Matrix cleaning is obviously more
difficult with higher packing densities or longer matrices. Higher
particle concentrations will require more frequent cleaning, and the
difficulty of cleaning is most probably related to particle size and
susceptibility. If parallel, cyclical systems are to be applied to fine
particle control, all of these factors will have a strong effect on the
duty cycle of the individual magnets and the design of the auxiliary
cleaning apparatus. The economic viability of the process could well
hinge on the successful demonstration of either a cyclical system with
reasonable duty cycles and simple mechanical construction, or preferably
a continuously operating system based on the carousel principle described
in Section 4 or some other innovative design.
102
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SECTION 7
ECONOMIC COMPARISON OF HGMS AND
CONVENTIONAL CONTROL TECHNOLOGY
BASIS OF ECONOMIC COMPARISON
To compare the capital cost and power requirements of alternate
methods of fine particle control, one must insure that the estimates are
based on equipment that will achieve comparable results in a specific
application. The design of particulate control equipment is usually
highly specific to the particular application. For example, the design
parameters of an electrostatic precipitator (ESP) for a coal-fired
boiler might be quite different from those for a basic oxygen furnace
because of differences in particulate and gas stream characteristics.
HGMS performance data are currently available only for the collection of
particulate that originated in basic oxygen and electric arc steel-
making furnaces. Estimates of capital cost and energy requirements
derived from these data should therefore be compared only to estimates
for conventional technology applied to BOF's or EAF's. Even for a
singular type of application, differences will certainly exist in the
actual cost and energy requirements of equipment designed for specific
furnaces. The estimates presented in this report are therefore intended
to serve as an approximate economic comparison of HGMS with conventional
control technology applied only to steel-making furnaces.
The preliminary nature of this analysis combined with the uncer-
tainty that would have been introduced by comparing different existing
industrial applications (many of which are retrofits) ruled out an
inquiry to the iron and steel industry to determine actual cost and
energy requirements of existing BOF and EAF particulate control equip-
ment. Instead the analysis was based on a study conducted by the
103
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Industrial Gas Cleaning Institute, (IGCI) for EPA in 1971 [88], The
published report of this study contains actual vendor quotations for
high-efficiency electrostatic precipitators and venturi scrubbers
designed to control particulate emissions from a typical basic oxygen
furnace. At the time the study was conducted there were no baghouses in
use on BOF's in the United States. Electric arc furnaces were not
specifically addressed in the IGCI report. Although the particulate
matter emitted from the two types of furnaces is similar, the cost and
energy requirements for comparable control efficiencies could be some-
what different.
The IGCI study reports flange-to-flange control device costs, total
installed capital costs, and operating costs in 1971 dollars. Since
information on HGMS auxiliary equipment (ducts, dampers, collected dust
handling equipment, etc.) and installation costs are not yet established
for particulate control applications, the comparison presented here is
limited to flange-to-flange device costs and operating power require-
ments. Also since HGMS collection of airborne particulate has been
demonstrated only on bench-scale and pilot-scale equipment, the scale-up
factor is very large. The full-scale HGMS estimates were based on
modules consisting of commercially available magnets (complete with
power supplies and cooling equipment) arranged in a multi-path parallel
flow configuration for cyclical operation. A 70-percent duty cycle
was assumed. A continuously operating system might offer better
utilization of magnetized volume and hence lower capital costs, but
the application of continuous units to gas streams has not yet been
demonstrated.
All equipment costs reported here have been escalated from their
original reported date of applicability to mid-1977 using the Marshall
and Swift equipment cost index [89].
COST AND ENERGY ESTIMATES FOR CONVENTIONAL TECHNOLOGY
The IGCI study considered several alternative designs. Electro-
static precipitators are installed only in open-hood BOF systems to
avoid explosion hazards. Wet scrubbers and, presumably, high gradient
104
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magnetic separators can function safely and effectively in either open-
hood or closed-hood systems. To allow better comparison, the ESP and
venturi scrubber with the most nearly comparable design features (open-
hood system, largest gas flows, equal collection efficiencies) were
selected. A summary of the ESP and, scrubber designs is presented in
Table 9. The 99.9 percent collection efficiency specified by the IGCI
is a very stringent requirement which in most cases will exceed the EPA
new source performance standards for basic oxygen furnaces [90]. The
designs reported by the IGCI may be optimistic in that higher pressure
drop requirements have been reported for wet scrubbers operating at
equivalent or lower collection efficiencies [91,92].
COST AND ENERGY ESTIMATES FOR HGMS
The mathematical model that was used to analyze the pilot-scale
HGMS experimental results also provides a quantitative approach to the
optimization of process operating conditions for a particular appli-
cation. The process cannot be completely optimized, however, until
additional information is developed on alternative magnet designs
oriented directly to gas-stream application and until a wider experi-
mental data base is obtained to further verify the model. The HGMS
units presently in service in the clay industry are basically of a
singular design: 2.13-m diameter iron-bound solenoids that require 400-
500 kW to produce applied fields on the order of 2 T over a pole gap
compatible with a 50-cm matrix. For fine particle collection in gas
streams, much lower fields may be sufficient, and the approach to cost-
optimized design could be significantly different. Also, because of the
large quantities of gas that must be cleaned, matrix pressure drop is a
more critical parameter that could also affect magnet design and matrix
configuration.
To conduct this analysis cost estimates were obtained from a com-
mercial vendor [93] for large iron-bound solenoids of the conventional de-
sign with modifications that would make them more compatible with large gas
flows and with the operating conditions proven effective in the pilot-
scale tests. The estimates are summarized in Table 10. Power requirements
105
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TABLE 9. DESIGN AND COST SUMMARY OF CONVENTIONAL BOF CONTROL EQUIPMENT
Control Device
System of Units
Process Gas Flow
Outlet Particulate Loading
Total Mass Collection Efficiency
Pressure Drop
Total Power Requirement
Flange-to-Flange Cost (1977 $)
Electrostatic Precipitator
SI
481 m3/s^
0.011 g/m
99.9%
- -
1525 kl/4)
English
1,020,000 ACFM^
0.0046 gr/ACF^
99.9%
- -
2045 hp^
$1,944,000
Wet Scrubber
SI
309 m3/s^2^
0.016 g/m3^2^
99.9%
10.5 kPa^3^
5605 kW^
English
655,000 ACFM^
0.0072 gr/ACF^
99.9%
42 in. H20^3^
7518 hp(5)
$640,000
(1) At precipitator outlet: 343° C (650° F), -1.5 kPa (-6 in. H20) gauge, 11 percent water by volume.
(2) At scrubber outlet: 82° C (180° F), -11.2 kPa (-45 in. H20) gauge, 57.4 percent water by volume.
(3) Across scrubber and mist eliminator.
(4) Report does not state explicitly what this includes. It definitely includes precipitator energization
and pressure drop, but may also include some duct pressure drop.
(5) Report does not state explicitly what this includes. It definitely includes liquid pumping and scrubber
pressure drop, but may also include some duct pressure drop.
-------
TABLE 10. COST ESTIMATES FOR LARGE IRON-BOUND SOLENOIDS APPLICABLE
TO FINE PARTICLE CONTROL
Matrix Diameter
Matrix Length
Field Capability
Cost, (1977)
3.15 m
15 cm
0.4 T
$320,000
3.15 m
30 cm
0.4T
$370,000
of these units were scaled from existing data on 2.13-m diameter by 50-
cm matrix length units. Several factors can influence magnet power
requirements, but with similarly designed solenoids to a good approxi-
mation the power requirement for a given field is directly proportional
to the diameter and to the length of the cylindrical magnetized volume.
Reported data for the 2.13-m units and estimates for the two larger
units are shown in Figure 49. References for the data are indicated on
the figure.
The operating temperature and gas volume of the I6CI electrostatic
orecipitator were taken as operating conditions for the HGMS. Theo-
retically higher temperatures will have an adverse effect on HGMS
performance by increasing gas viscosity and decreasing particle magnetic
susceptibility. A correction was made for the viscosity difference but
the temperature dependence of particle susceptibility has not yet been
determined. The design was also based on basic oxygen furnace dust
since that was the basis of the IGCI designs. The electric arc furnace
dust used in the pilot-scale tests would require a more expensive design
in terms of capital investment and/or power requirements because of the
lower magnetic susceptibility of the dust. This does not necessarily
mean that all electric arc furnaces would be more expensive to control
or that the design developed here would be appropriate for all BOFs.
Since the composition of the dust varies from furnace to furnace, the
magnetic susceptibility most probably does also.
Since the IGCI designs stipulated 99.9 percent collection efficiency
without stating the particle size distribution, the HGMS design was
107
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1000 i
900
800
700
600
500
400
300
200
85 100
O
0_
O
LJ
or
o
LJ
or
90
80
70
60
40
30
20
10
01
x 15cm
0.2 0.3 0.4 Q5 0.6 0.8 1.0
APPLIED FIELD, T
2.0
3.0
Figure 49. Reported and estimated power requirements for
iron-bound solenoids of various configuration.
108
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based on a reasonably comparable specification of 99.9 percent collection
of 1.0 ym BOF particles. The procedure for estimating flange-to-flange
costs and power requirements is outlined below.
(1) Basis: 99.9 percent collection of 1.0 ym particles
(2) Design equation:
EFLR^
(34)
P = exp
(3) Determination of operating parameters: Seven of the 3.15-m
o
diameter units will handle the 481 m /s gas flow with a
superficial gas velocity of 8.8 m/s. With BOF dust and the
same matrix material used in the pilot-scale experiments,
E = 0.09 and a = 25 ym. For either F * 0.005, L = 30 cm or
F = 0.010, L = 15 cm, the field requirement is determined
as follows.
particle diameter 1.0 ym
required collection efficiency 99.9%
from Equation (34), R = 1.27
\*
from Equation (18), K = 2.94
log K = 0.47
from Figure 50, log W = -0.64
VI = 0.229
from Equation (16), HQ = 3.5 x 105 A/m
B = 0.44 T
o
For additional reference, with these operating parameters the
calculated collection efficiency of 0.5 ym particles is
98.7 percent.
(4) Calculation of power requirements:
(a) With a shorter, more dense matrix (L = 15 cm, F = 0.010):
From Figure 49 power per magnet is 27.5 kW for a total
magnet power requirement of 192 kW. From Equation (35)
109
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5 -
log W
-2h
A=l.5
G = 0
-2
-I
0
log K
Figure 50. Lines of constant capture radius with inertial forces included
and A=l.5.
no
-------
corrected for lower gas density at 343° C, matrix pressure
drop is 1.1 kPa (4.4 in. H20). With 60 percent fan
efficiency, fan power is 885 kW. Total power requirement
is thus 1077 kW.
(b) With a longer, less dense matrix (L = 30 cm, F = 0.005):
From Figure 49 power per magnet is 55 kW for a total
magnet power requirement of 385 kW. From Equation (35)
corrected for lower gas density at 343° C matrix pressure
drop is 0.9 kPa (3.5 in. H20). With 60 percent efficiency,
fan power in 702 kW. Total power requirement is thus
1087 kW.
(5) Calculation of capital cost: Since the power requirements of
the two alternatives considered are virtually identical, the
shorter, more dense matrix is clearly superior because of the
lower cost of the magnets. With a 70 percent duty cycle, a
total of ten magnets would be required for a total flange-to-
flange cost of $3,200,000.
COMPARISON OF HGMS AND CONVENTIONAL TECHNOLOGY
The estimated cost and energy requirements of the HGMS device are
compared with conventional technology in Table 11. Since the assumed
gas temperature, pressure, and water content were not the same in the
wet scrubber as in the electrostatic precipitator and HGMS, normal-
ization of costs and energy requirements by the actual gas flow in each
device does not provide a direct comparison. To refer the estimates to
an equivalent application the scrubber gas volume was corrected to the
operating gas conditions of the precipitator and HGMS.
Table 11 indicates that on a flange-to-flange basis the scrubber
has the lowest capital cost. This may be misleading, however, since the
scrubber would require an associated water treatment system not required
by either the precipitator or HGMS. With the design utilized for these
estimates the capital cost of the HGMS is somewhat higher than the pre-
cipitator but still well within the economic ballpark considering the
accuracy of the estimates. The estimates of power requirement show that
the scrubber requires several times more energy than either the pre-
cipitator or the HGMS to do a comparable job. The HGMS is estimated to
111
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TABLE 11. COMPARISON OF H6MS AND CONVENTIONAL TECHNOLOGY
Device
Collection Efficiency, %
3
Flange-to-Flange Cost, $/m /s
$/CFM
Power Requirement, kW/m /s
hp/1000 CFM
ESP
99.9
4042
1.91
3.2
2.0
Scrubber
99.9
2762
1.30
24.2
15.3
HGMS
99.9
6653
3.14
2.2
1.4
All estimates are referred to clean gas at 343° C (650° F), -1.5 kPa
(-6 in. HpO) gauge, 11 percent water by volume.
be more efficient than the precipitator in terms of energy utilization.
Again, however, the accuracy of the estimates must be considered. Also
the precipitator estimate may include some unintended contribution for
duct pressure drop.
Several general conclusions about optimization of the HGMS process
can be drawn from the above results. First, it would appear that matrix
length should be kept short at the expense of other variables to achieve
the most cost effective designs. The limiting factors of this situation
might be good flow distribution and homogeneity of the applied magnetic
field, both of which become more difficult to achieve as the aspect ratio
of the matrix increases.
The relative importance of capital and operating costs can vary
with individual applications according to the company's amoritization
methods, tax situation and so forth. Compared to the precipitator
system, however, the HGMS estimates of Table 11 imply a need to lower
capital costs perhaps at the expense of power requirement if necessary.
One way to accomplish this reduction would be to increase the superficial
gas velocity and thus reduce the required number of magnets. Since the
fan power requirement is already considerably higher than the magnet power
requirement, this increase in velocity might be accompanied by a reduction
112
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in matrix packing density to lower pressure drop and an increase in the
applied field of each magnet to maintain the high collection efficiency.
Another approach would be to cool the gas further before cleaning it.
The volume of gas cleaned could be reduced allowing a smaller number of
magnets to handle the flow. This savings would of course be at least
partially offset by the investment required to accomplish the additional
gas cooling.
While these adjustments could in principle lead to a more optimum
design, they also point to a potential problem. The loading time of the
collection matrix (time between cleanings) is apparently directly pro-
portional to the mass of the matrix and inversely proportional to the
gas velocity and to the particulate concentration. Other factors such
as applied field and particle susceptibility may also be important. With
a matrix density of 0.010, a matrix length of 15 cm, and a superficial
velocity of 8.8 m/s (the design used to obtain the economic estimates
3 3
above), a matrix cleaning a dirty gas concentration of 10 g/m (4.4 gr/ft )
would collect its own mass of dust in 2.2 minutes. By the time the matrix
collected its own mass, increased pressure drop or decreased collection
efficiency would probably dictate that the matrix be cleaned. To
achieve a 70 percent duty cycle would then require that the matrix be
cleaned and returned to service in slightly under 1 minute. In batch-
wise operation the gas flow and magnetic field would be temporarily shut
off to allow in-situ"matrix cleaning. The large iron-bound solenoids in
commercial use are reported to have a field decay time on the order of
60-90 seconds although technology has now been developed that accelerates
this decay time to 5 to 10 seconds [5]. Damper systems to accomplish
the necessary flow switching in this relatively short time frame could be
relatively costly and subject to maintenance problems. The indicated
optimization technique of increasing gas velocity and decreasing matrix
density would shorten the loading time and the cleaning time required to
maintain a 70 percent duty cycle. The reduction of gas temperature
would also shorten the loading and required cleaning times by creating a
more concentrated dust stream.
113
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This analysis points to the need of additional developmental work
in the design of magnets, matrices, and cleaning systems specifically
geared to application in high velocity gas streams with moderate and
high dust concentrations. The short estimated cycle times provide a
strong incentive to develop continuously operating rather than cyclic
systems.
The comparison of alternate BOF emission control methods suggests
that HGMS is a viable alternative for fine particle control in systems
where the particulate has a relatively high magnetic susceptibility. In
addition to reasonable projected capital cost and energy requirements
the process offers other potential advantages over conventional technology.
The high throughput suggests smaller space requirements than electrostatic
precipitators or baghouses. This could be important in any application
but especially perhaps in a retrofit situation. Stainless steel wool is
a common, relatively inexpensive material which is resistant to high
temperature and corrosive environments. The lack of any sparking
mechanism should make the process amenable to dry, combustible environ-
ments with proper design precautions. Unlike many wet scrubber systems
the HGMS process would not turn an air pollution problem into a water
pollution problem. Even if water should be employed to assist in matrix
cleaning the quantity would be much less than in scrubber systems.
These attractive features of the HGMS process strongly support an
argument for the continued development of the technology with early
emphasis on identifying the most appropriate potential applications
within the iron and steel industry as well as the ferroalloy industry.
A program should be undertaken to demonstrate the fine particle control
capability of HGMS with a pilot-scale device operating on a slipstream
of one or more actual industrial processes. Continued research should
also be supported to develop and test HGMS system designs specifically
oriented toward gas stream applications so that capital costs and power
requirements can be minimized and efficient, practical matrix cleaning
systems can be demonstrated.
114
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SECTION 9
REFERENCES
1. Oberteuffer, J.A., Magnetic Separation: A Review of Principles,
Devices and Applications, IEEE Transactions on Magnetics, Vol. MAG-
10, No. 2, June 1974, pp. 223-238.
2. Kolm, H., J. Oberteuffer, and D. Kelland, High Gradient Magnetic
Separation, Scientific American, Vol. 233, No. 5, November 1975,
pp. 47-53.
3. Luborsky, F.E., High Field-High Gradient Magnetic Separation: A
Review, Magnetism and Magnetic Materials—1975, 21st Annual Confer-
ence, Philadelphia, America'n Institute of Physics Conference Proceedings
1976.
4. Oder, R.R., High Gradient Magnetic Separation Theory and Applications,
IEEE Transactions on Magnetics, Vol. MAG-12, No. 5, September 1976,
pp. 428-435.
5. lannicelli, J., New Developments in Magnetic Separation, IEEE
Transactions on Magnetics, Vol. MAG-12, No. 5, September 1976, pp.
436-443.
6. Franz, S.G., Magnetic Separator, U.S. Patent 2,074,085, March 16,
1937.
7. Jones, G.H., Magnetic Separators, U.S. Patent 3,346,116, October
10, 1967.
8. Oder, R.R., and C.R. Price, Brightness Beneficiation of Kaolin
Clays by Magnetic Treatment, The Journal of the Technical Association
of the Pulp and Paper Industry, Vol. 56, No. 10, October 1973, pp.
75-78.
9. Oder, R.R., The Magnetic Processing of Kaolin Clay, Pulp and Paper
Magazine of Canada, Vol. 75, No. 10, October 1974, pp. T366-T369.
10. lannicelli, J., Development of High Extraction Magnetic Filtration
by the Kaolin Industry of Georgia, IEEE Transactions on Magnetics,
Vol. MAG-12, No. 5, September 1976, p. 489.
11. lammartino, N.R., New Tasks for Magnetism, Chemical Engineering,
Vol. 81, No. 1, January 7, 1974, pp. 50-52.
12. Kelland, D.R., High Gradient Magnetic Separation Applied to Mineral
Beneficiation, IEEE Transactions on Magnetics, Vol. MAG-9, No. 3,
September 1973, pp. 307-310.
115
-------
13. Lawver, J.E., and D.M. Hopstock, Wet Magnetic Separation of Weakly
Magnetic Minerals, Minerals Science Engineering, Vol. 6, No. 3,
July 1974, pp. 154-172.
14. Kelland, D.R., and E. Maxwell, Oxidized Taconite Beneficiation by
Continuous High Gradient Magnetic Separation, IEEE Transactions on
Magnetics, Vol. MAG-11, No. 5, September 1975, pp. 1582-1584.
15. Vives, D.L., L.J. Hirth, and W.H. Summer!in, Direct Reduction and
Magnetic Beneficiation of Alabama Brown Ore with Lignite, IEEE
Transactions on Magnetics, Vol. MAG-12, No. 5, September 1976,
pp.,490-492.
16. Murray, H.H., Beneficiation of Selected Industrial Minerals and
Coal by High Intensity Magnetic Separation, IEEE Transactions on
Magnetics, Vol. MAG-12, September 1976, pp. 498-502.
17. Ergun, S., and E.H. Bean, Magnetic Separation of Pyrite from Coals,
Report of Investigation No. 7181, U.S. Bureau of Mines, Pittsburgh,
PA., September 1968.
18. Trindade, S.C., and H.H. Kolm, Magnetic Desulfurization of Coal, IEEE
Transactions on Magnetics, Vol. MAG-9, No. 3, September 1973,
pp. 310-313.
19. Oder, R.R., Pyritic Sulfur Removal from Coals by High Gradient
Magnetic Separation, Proceedings of Second National Conference on
Complete WateReuse, 1975, pp. 315-322.
20. Cohen, E. and J.A. Good, The Application of a Superconducting
Magnet System to the Cleaning and Desulfurization of Coal,
IEEE Transactions on Magnetics, Vol. MAG-12, No. 5, September 1976,
pp. 503-506.
21. Maxwell, E., D.R. Kelland, and I.Y. Akoto, High Gradient Magnetic
Separation of Mineral Particulates from Solvent Refined Coal, IEEE
Transactions on Magnetics, Vol. MAG-12, No. 5, September 1976,
pp. 507-510.
22. Lin, C.J., Y.A. Liu, D.L. Vives, M.J. Oak, G.E. Crow, and E.L. Huffman,
Pilot-Scale Studies of Sulfur and Ash Removal from Coals by High
Gradient Magnetic Separation, IEEE Transactions on Magnetics, Vol
MA6-12, No. 5, September 1976, pp. 513-521.
23. Monteiro, J.L.F., M. Saddy, and S.C. Trindade, Studies on Sulfur
Recovery from Coal Wastes and Prospective Magnetic Separation
Applications, IEEE Transactions on Magnetics, Vol. MAG-12, No. 5,
September 1976, p. 522.
116
-------
24. Fine, H.A., M. Lowry, L.F. Power, and G.H. Geiger, A Proposed Process
for the Desulfurization of Finely Divided Coal by Flash Roasting and
Magnetic Separation, IEEE Transactions on Magnetics, Vol. MAG-12,
No. 5, September 1976, pp. 523-527.
25. Oder, R.R., Magnetic Desulfurization of Liquefied Coals: Conceptual
Process Design and Cost Estimation, IEEE Transactions on Magnetics,
Vol. MAG-12, No. 5, September 1976, pp. 532-537.
26. Liu, Y.A., and C.J. Lin, Assessment of Sulfur and Ash Removal from
Coals by Magnetic Separation, IEEE Transactions on Magnetics, Vol.
MAG-12, No. 5, September 1976, pp. 538-550.
27. Maxwell, E., D.R. Kelland, and I.Y. Akoto, Magnetic Separation of
Solids from Liquified Coal, IEEE Transactions on Magnetics, Vol.
MAG-12, No. 6, November 1976, p. 907.
28. Kazarian, R., Researchers Developing Method to Remove Impurities
From-Coal, National Science Foundation News, NSF PR77-50, May 10,
1977, pp. 1-3.
29. Marston, P.G., and J.A. Oberteuffer, The Application of High Gradient
Magnetic Separation (HGMS) to the Treatment of Steel Industry Waste
Waters, Second International Congress on Industrial Wastes, Stockholm,
Sweden, February 1975.
30. Oberteuffer, J.A., I. Wechsler, P.G. Marston and M.J. McNallan,
High Gradient Magnetic Filtration of Steel Mill Process and Waste
Waters, IEEE Transactions on Magnetics, Vol. MAG-11, No. 5,
September 1975, pp. 1591-1593.
31. Oberteuffer, J.A., Engineering Development of High Gradient Magnetic
Separators, IEEE Transactions on Magnetics, Vol. MAG-12, No. 5,
September 1976, pp. 444-449.
32. Harland, J.R., L. Nilsson and M. Wallin, Pilot-Scale High Gradient
Filtration of Steel Mill Wastewater, IEEE Transactions, Vol. MAG-12,
No. 6, November 1976, pp. 904-906.
33. deLatour, C., Magnetic Separation in Water Pollution Control, IEEE
Transactions on Magnetics, Vol. MAG-9, No. 3, September 1973, pp. 314-
316.
34. Oder, R.R., and B.I. Horst, Wastewater Processing with High Gradient
Magnetic Separators, Proceedings of Second National Conference on
Complete WateReuse, 1975, pp. 887-897.
35. deLatour, C. and H. Kolm, Magnetic Separation in Water Pollution
Control-II, IEEE Transactions on Magnetics, Vol. MAG-11, No. 5,
September 1975, pp. 1570-1572.
117
-------
36. Ricci, L.J., Antipollution Processes Top Boston AIChE Meeting,
Chemical Engineering, Vol. 82, No. 21, October 13, 1975, pp. 73-75.
37. Petrakis, L. and P.P. Ahnen, Use of High Gradient Magnetic Separation
Techniques for the Removal of Oil and Solids from Water Effluents,
IEEE Transactions on Magnetics, Vol. MAG-12, No. 5, September 1976,
pp. 486-488.
38. Yadidia, R., A. Abeliovich, and G. Belfort, Algae Removal by High
Gradient Magnetic Filtration, Environmental Science and Technology
Vol. 11, No. 9, September 1977, pp. 913-916.
39. Melville, D., F. Paul, and S. Roath, High Gradient Magnetic Separation
of Red Cells from Whole Blood, IEEE Transactions on Magnetics, Vol.
MAG-11, No. 6, November 1975, pp. 1701-1704.
40. Kolm, H.H., Research Needs In Magnetic Separation, IEEE Transactions
on Magnetics, Vol. MAG-12, No. 5, September 1976, pp. 450-454.
41. Kolm, H.H., The Large-Scale Manipulation of Small Particles, IEEE
Transactions on Magnetics, Vol. MAG-11, No. 5, September 1975, pp.
1567-1569.
42. Ranz, W.E. and J.B. Wong, Impaction of Dust and Smoke Particles on
Surface and Body Collectors, Industrial and Engineering Chemistry,
Vol. 44, No. 6, June 1952, pp. 1371-1381.
43. Davies, C.N., The Separation of Airborne Dust and Particles, Proceedings
of the Institute of Mechanical Engineering, Vol. 18, 1952, pp. 185-
198.
44. Fuchs, N.A., The Mechanics of Aerosols. Pergamon Press, New York,
1964.
45. Strauss, W., Industrial Gas Cleaning, Pergamon Press, New York,
1966.
46. Loffler, F., Collection of Particles by Fiber Filters, in Air Pollution
Control, Part 1, edited by W. Strauss, Wiley-Interscience, New
York, 1971.
47. Bean, C.P., Theory of Magnetic Filtration, Bulletin of American
Physical Society, Vol. 16, 1971, p. 350. (abstract only)
48. Watson, J.H.P., Magnetic Filtration, Journal of Applied Physics, Vol
44, No. 9, September 1973, pp. 4209-4213.
49. Oberteuffer, J.A., High Gradient Magnetic Separation, IEEE Transactions
on Magnetics, Vol. MAG-9, No. 3, September 1973, pp. 303-306.
118
-------
50. Oder, R.R. and C.R. Price, HGMS: Mathematical Modeling of Commercial
Practice, Magnetism and Magnetic Materials-1975, 21st Annual Confer-
ence, Philadelphia, American Institute of Physics Conference Proceedings,
1976. a-
51. Watson, J.H.P., Theory of Capture of Particles in Magnetic High-
Intensity Filters, IEEE Transactions on Magnetics, Vol. MAG-11,
No. 5, September 1975, pp. 1597-1599.
52. Cowen, C., F.J. Friedlaender, and R. Jaluria, High Gradient Magnetic
Field Particle Capture on a Single Wire, IEEE Transactions on Magnetics,
Vol. MAG-11, No. 5, September 1975, pp. 1600-1602.
53. Luborsky, F.E. and B.J. Drummond, High Gradient Magnetic Separation:
Theory Versus Experiment, IEEE Transactions on Magnetics, Vol. MAG-
11, No. 6, November 1975, pp. 1696-1700.
54. Cummings, D.L., D.A. Himmelblau, J.A. Oberteuffer, and G.J. Powers,
Capture of Small Paramagnetic Particles by Magnetic Forces from Low
Speed Fluid Flows, AIChE Journal, Vol. 22, No. 3, May 1976, pp. 569-
575.
55. Aharoni, A., Traction Force on Paramagnetic Particles in Magnetic
Separators, IEEE Transactions on Magnetics, Vol. MAG-12, No. 3,
May 1976, pp. 234-235.
56. Luborsky, F.E., and B.<3. Drummond, Buildup of Particles on Fibers in
a High Field-High Gradient Separator, IEEE Transactions on Magnetics,
Vol. MAG-12, No. 5, September 1976, pp. 463-465.
57. Cowen, C., F.J. Friedlaender, and R. Jaluria, Single Wire Model of
High Gradient Magnetic Separation Processes I, IEEE Transactions on
Magnetics, Vol. MAG-12, No. 5, September 1976, pp. 466-470.
58. Cummings, D.L., D.C. Prieve, and G.J. Powers, The Motion of Small
Paramagnetic Particles in a High Gradient Magnetic Separator, IEEE
Transactions on Magnetics, Vol. MAG-12, No. 5, September 1976,
pp. 471-473.
59. Stekly, Z.J.J. and J.V. Minervini, Shape Effect of the Matrix on the
Capture Cross Section of Particles in High Gradient Magnetic Separation,
IEEE Transactions on Magnetics, Vol. MAG-12, No. 5, September 1976,
pp. 474-479.
60. Birss, R.R., R. Gerber, and M.R. Parker, Theory and Design of Axially
Ordered Filters for High Intensity Magnetic Separation, IEEE Transactions
on Magnetics, Vol. MAG-12, No. 6, November 1976, pp. 892-894.
119
-------
61. Uchiyama, S., S. Kondo, M. Takayasu, and I. Eguchi, Performance of
Parallel Stream Type Magnetic Filter for HGMS, IEEE Transactions
of Magnetics, Vol. MAG-12, No. 6, November 1976, pp. 895-897.
62. Cowen C., F.J. Friedlaender, and R. Jaluria, Single Wire Model of
High Gradient Magnetic Separation Processes II, IEEE Transactions
on Magnetics, Vol. MAG-12, No. 6, November 1976, pp. 898-900.
63. Clarkson, C.J., D. Kelland, and T.B. King, Model for Calculation of
Capture Radii of High Gradient Magnetic Separator at Moderate
Reynolds Numbers, IEEE Transactions on Magnetics, Vol. MAG-12,
No. 6, November 1976, pp. 901-903.
64. Lawson, W.F., The Dynamics of Paramagnetic Particles Near a
Magnetized Wire, M.S. Thesis, Department of Physics, West Virginia
University, Morgantown, WV, 1976. (unpublished).
65. Lawson, W.F., W.H. Simons, and R.P. Treat, The Dynamics of a
Particle Attracted by a Magnetized Wire, Journal of Applied
Physics, Vol. 48, No. 8, August 1977, pp. 3213-3224.
66. Zebel, G., Deposition of Aerosol Flowing Past a Cylindrical Fiber
in a Uniform Electric Field, Journal of Colloid Science, 20,
1965, pp. 522-543.
67. Krupp, H., Particle Adhesion, Advances in Colloid and Interface
Science, Vol. 1, 1967, pp. 111-239.
68. Ketkar, A., and D.V. Keller, Adhesion of Micron-Sized Limestone
Particles to a Massive Coal Substrate, Journal of Adhesion,
Vol. 7, 1975, pp. 235-251.
69. Visser, J., On Hamaker Constants: A Comparison between Hamaker
Constants and Lifshitz - Van der Waals Constants, Advances in
Colloid and Interface Science, Vol. 3, 1972, pp. 331-363.
70. Marks' Standard Handbook for Mechanical Engineers, T. Baumeister,
Editor-in-Chief, 7th edition, McGraw-Hill Book Company, New York, 1967.
71. Katari, V., G. Issacs and T.W. Devitt, Trace Pollutant Emissions
from the Processing of Metallic Ores, EPA-650/2-74-115, October
1974.
72. Hedley, W.H., S.M. Mehta, C.M. Moscowitz, A.D. Snyder, H.H.S. Yu,
and D.L. Zanders, Sources and Characterization of Fine Particular
Test Dusts, EPA-650/2-74-117, November 1974. " ~
120
-------
73. Dealy, J.O., and A.M. Killin, Engineering and Cost Study of the
Ferroalloy Industry, EPA-450/2-74-008, May 1974.
74. Sears, F.W., and M.W. Zemansky, University Physics, Part 2, 3rd
edition. Addison-Wesley Publishing Company, Inc., Reading,
MA, 1964.
75. Oberteuffer, J.A. and M. Golay, A Review and Analytical Evalu-
ation of the Application of High Gradient Magnetic Separation
to Fine Particle Scrubbing, report submitted by Magnetic Engi-
neering Associates under subcontract to Research Triangle Insti-
tute, December, 1975.
76. Sala Magnetics, Inc. Cambridge, MA, SMI Bulletin No. D052401-
7527GB/775.
77. Guichard, J.C., Aerosol Generation Using Fluidized Beds, in Fine
Particles-Aerosol Generation, Measurement, Sampling and Analysis,
edited by B.Y.H. Liu, Academic Press, New York, 1976, pp.173-194.
78. Moreno, F. and D. Blann, Large Flow Rate Redispersion Aerosol
Generator, in Fine Particles-Aerosol Generation. Measurement
Sampling and Analysis, edited by B.Y.H. Liu, Academic Press, New
York, 1976, pp. 195-218.
79. Willeke, S., C. Lo, and K.T. Whitby, Dispersion Characteristics of
a Fluidized Bed, Aerosol Science, Vol. 5, 1974, pp. 449-455.
80. Ragland, J.W., D.H. Pontius, and W.B. Smith, Design, Construct,
and Test a Field Usable Prototype System for Sizing Particles
Smaller than 0.5 m Diameter , Special summary report submitted
by Southern Research Institute to EPA under Contract No. 69-
02-2114, Task 8, EPA Task Officer Bruce Harris, February 1976,
(unpublished).
81. Blann, D.D., K.A. Green, and L.W. Anderson, Design, Fabrication,
and Installation of a Particulate Aerodynamic Test Facility,
EPA-650/2-74-103, October 1974.
82. Gotterba, J., Characterization and Optimization of the EPA
High Mass Flowrate Aerosol Generator, Task report submitted
by Aerotherm Corporation to EPA under Contract No. 68-02-1318,
Task 28, EPA Task Officer Dale Harmon, July 1976, (unpublished).
83. Roulliard, E.E.A., Experimental Errors in Sampling Dust Laden
Gas Streams, South African Council for Scientific and Industrial
Research, Chemical Engineering Group, CSIR Special Report File
No. 66/51/4510/2, Pretoria, December 1971.
121
-------
84- Harris, D.B., Procedures for Cascade Impactor Calibration and
Operation in Process Streams, EPA-600/2-77-004, January 1977.
85. Gushing, K.M., G.E. Lacey, J.D. McCain, and W.B. Smith, Particulate
Sizing Techniques for Control Device Evaluations Cascade Impactor
Calibrations. EPA-600/2-76-280, October 1976.
86. Hay, R.D., Magnetic Engineering Associates, Inc., Cambridge,
MA, Personal communication, April 29, 1976.
87. Oder, R.R., Bechtel Corporation, San Francisco, CA, Personal
communication, January 6, 1977.
88. Hardison, L.C. and C.A. Greathouse, Air Pollution Control
Technology and Costs in Nine Selected Areas, final report of
Industrial Gas Cleaning Institute to EPA under Contract No.
68-02-0301, National Technical Information Service, Springfield,
Virginia, PB-222 746, September 1972.
89. Economic Indicators, Chemical Engineering, Vol. 84, No. 19,
September 12, 1977, p. 7.
90. Federal Register, March 8, 1974, Vol. 39, p. 9308.
91. Steiner, B.A., Air-Pollution Control in the Iron and Steel Industry,
International Metals Review, September 1976.
92. Delaney, E. L., Development Document for Effluent Limitations
Guidelines and New Source Performance Standards for the Steel
Making Segment of the Iron and Steel Manufacturing Point Source
Category. EPA-440/l-74-024a, June 1974.
93. Wechsler, I. Sala Magnetics, Inc., Cambridge, MA, Personal
communication, October 18, 1977.
122
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SECTION 10
BIBLIOGRAPHY
1. Allen, J.W., Magnetic Separator, U.S. Patent 3,819,515, June 25,
1974.
2. Aubrey, W.M., J.M. Karpinski, and D.S. Cahn, Magnetic Separator
Method and Apparatus, U.S. Patent 3,608,718, September 28, 1971.
3. Davis, C.W.V., Magnetic Separator, U.S. Patent 3,062,376,
November 6, 1962.
4. Fraas, F., Alternating Field Magnetic Separator, U.S. Patent
3,988,240, October 26, 1976.
5. Fraas, F., The Matrix-Type Magnetic Separator, Report of Investi-
gation No. 6722, U.S. Bureau of Mines, Pittsburgh, Pennsylvania,
1966.
6. Kolm, H.H., Magnetic Device, U.S. Patent 3,567,026, March 2,
1971.
7. Kolm, H.H., Process for Magnetic Separation, U.S. Patent
3,676,337, July 11, 1972.
8. Marston, P.G., Magnet Design and Its Effect on the Economics
of High Gradient Magnetic Separation Processes, Proceedings
of the High Gradient Magnetic Separation Symposium, Massachusetts
Institute of Technology, Cambridge, Mass., May 1973. pp. 25-37.
9. Marston, P.G., I. Wechsler, and J.J. Nolan, Magnetic Separation
Method, U.S. Patent 3,887,457, June 3, 1975.
10. Marston, P.G., and J.J. Nolan, Moving Matrix Magnetic Separator,
U.S. Patent 3,920,543, November 18, 1975.
11. Marston, P.G., J.J. Nolan, and L.M. Lontai, Magnetic Separator
and Magnetic Separation Method, U.S. Patent 3,627,678,
December 14, 1971.
12. Nolan, J.J. and P.G. Marston, Flow Control Unit for Magnetic
Matrix, U.S. Patent 4,025,432, May 24, 1977.
123
-------
13. Nolan, J.J., P.G. Marston, and L.M. Lontai, Multiple Matrix Magnetic
Separation Device and Method, U.S. Patent 3,770,629, November 6,
1973.
14. Oder, R.R., Method for Magnetic Beneficiation of Particle Dispersions,
U.S. Patent 3,985,646, October 12, 1976.
15. Sakata, S., A. Yoshikawa, and A. Tasaki, Magnetic Separation of
Aerosol Particles from Air Flow, Japan Journal of Applied Physics,
Vol. 15, No. 10, 1976, pp. 2017-2018.
16. Watson, J.H.P. and D. Hocking, The Beneficiation of Clays Using
a Superconducting Magnetic Separator, IEEE Transactions on
Magnetics, Vol. MAG-11, No. 5., September 1975, pp. 1588-1590.
124
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APPENDIX
DERIVATION OF THE EQUATIONS OF MOTION FOR A PARAMAGNETIC
PARTICLE IN A NONUNIFORM MA6NETOSTATIC FIELD
The energy of a magnetostatic field established in a linear,
homogeneous, and isotropic medium where B = y H is given by
where TQ = energy, joule
HQ = magnetic field strength, ampere/meter
Bn = magnetic flux density, tesla
-7
UQ = permeability of free space, 4Tr x 10 henry/meter.
dv = differential volume element.
The indicated integration is extended over all space. If a paramagnetic,
spherical particle of volume V-| (meter ) and radius R (meter) is placed
in the magnetic field, the potential energy is given by
(J = i- / M • B dv
m 2 \ "
where Um = potential energy, joule
M = induced magnetization, ampere/meter.
The integration is over the volume of the particle, and within this
volume
8 " PQ(H + M). (A-3)
125
-------
The magnetic force F (newton) acting upon the paramagnetic
particle is given by
-> -> -. -*• r •+ •+
F = - VU = j V / M • B dv (A-4)
vi
where in cylindrical coordinates
Ir
Vl
Interchanging v and / gives
Ti
->• , /*"*""*" "*"
Fm = ^ J 7(M • BQ)dv (A-6)
According to equation (A-3)
B = MQ(H 4- M) = yH r
-------
In the linear, homogeneous, and is^trbpic medium
-> ->•
Bo = yoHo r>>R (A-ll)
and equation (A-6) then becomes
Fm =
Vl
= 1 XP0/V(H • HQ)dv.
The dielectric constant, y, and the susceptibility, x» are properties of
the particle, which is also as1»"umed"to be linear, isotropic, and homogeneous.
If the value of H is assumed to be relatively constant throughout the
volume V-j and approximately equal to H0i the integration can be performed
to yield
+ -». +
= - «•• "'u • HQ) / dv
.'/
and finally,
The magnetic field of interest is a combination of the uniform
background field H and the nonuniform field in the vicinity of a
ferromagnetic wire. The geometric arrangement shown in Figure A-l
will be used to obtain the appropriate expression for the magnetic
field strength.
127
-------
Figure A-l. Coordinate system for HGMS single-fiber model
The radius of the cylindrical ferromagnetic wire is a (meter), and the
radius of the spherical particle is R (meter). For r»a the velocity
of the particle is V and parallel to but in the negative x direction.
For r»a, the magnetic field strength is H and parallel to and in the
positive x direction.
Since the magnetic field strength is a vector, we may define a
magnetic potential $ such that
H = -V *.
(A-15)
128
-------
As stated before for a linear, isotropic, and homogeneous medium,
B = uH. (A-16)
The magnetic flux density B is solenoidal, hence
-*• -»•
V • B = 0. (A-17)
Introducing Equations (A-15) and (A-16) into (A-17) yields
•-•-•••
V - yH = -yV • V « * -uV * = 0, (A-18)
where in cylindrical coordinates
3r2 r 3r r
2
Due to symmetry, -5-E = 0; and we are left with
3r
+ + _
3r2 r 3r r2 362
1 It + I_1« . o. (A-20)
129
-------
A solution of the form
- rnCn(e)
(A-21)
can be shown to satisfy the partial differential equation where Cn(e)
are harmonic functions given by
Cn
Co
Cl
C2
C3
r"nC (0) rnCos (0)
n ii
log r
r'Xcos
r"2(A1Cos
r"3(A1Cos
0 + B,Sin 0)
20 + B-jSin 20)
0 + B^in 30)
1
r(A2Cos 0 + E
r2(A2Cos 20 H
•5
r (A2Cos 30 J
5pSin 0)
h B2Sin 20)
i- B2Sin 30)
where A-|, A2, B-j, and B2 are constants.
Inside the wire the magnetic potential is -Mr, e). If only first
order harmonic functions are considered and the origin is chosen as zero
potential where $-](0,0) = 0, then the r terms must disappear and the
potential is given by
$-|(r, 0) = r(A Cos 0 + B Sin 0)
(A-22)
and
H-|(r, 0) =_(A Cos 0 + B Sin e)r - (-A Sin 0 + B Cos 0)0. (A-23)
130
-------
At the boundary of the wire the tangential component H of H is zero for
6=0. Evaluating (a,0) from Equation (A-23) yields
H-,j6(a,0) =-B = 0. (A-24)
Therefore Equation (A-22) becomes
^(r.g) = rA Cos 9. (A-25)
Outside the wire the first order approximation for the magnetic
potential is given by
$2(r,9) = r-1(C Cos 9 + D Sin 9) + r (E Cos 9 + F Sin 9) (A-26)
and
H (r,9) = -F- ^2 (C Cos 9 + D Sin 9) + (E Cos e +F Sin e)l r
r (A-27)
_| r'^-C Sin 9 + D Cos 9) + (-E Sin 9 + F Cos 9)1 9.
Outside the wire where r»a
H2(r,0) - HQr. (A-28)
Therefore,
,0) = HQ r»a
(A-29)
Ho ,(r,0) = 0 r»a. (A"3°)
131
-------
Evaluating H9 ^(r.O) and H~ Q(r,0) from Equation (A-27) yields
£» I £» 0
H2 r(r,0) = H-^-- E = HQ r»a. (A-31)
.'. E «-H0. (A-32)
Also
H9 .(r,0) = -- F = 0 r»a. (A-33)
t. 5 0 I
By choosing F = D = 0, the boundary condition H0 Q(r,0) = 0 is satisfied
£»°
and Equation (A-26) becomes
$2(r,e) = r C Cos e - rHQ Cos e. (A-34)
Other information necessary to specify A in Equation (A-25) and C
in Equation (A-34) is obtained from analysis of the conditions at the
boundary of the ferromagnetic wire. Throughout the ferromagnetic wire
the magnetization is parallel and in the positive x direction. Thus the
tangential component Mt and normal component Mn are related to M by
M = Mnr + Mt e (A-35)
Mn = |M| Cos Q (A-36)
Mt = |M| Sin e. (A-37)
132
-------
It can be easily shown that the normal components of the magnetic flux
density across the boundary of the wire must be equal or
Bnl = Bn2
Inside the wire
nl
3r
r=a
(A-38)
(A-39)
Outside the wire
B = yh
n2 on2 '
r=a
(A-40)
In addition, the tangential components of the magnetic field strength
must be continuous and equal across the boundary of the wire or
Inside the wire
Outside the wire
Htl = Ht2
i
Htl = Hl,e = " r 98
r=a
t2 = H2,e = " r 39
r=a
(A-41)
(A-42)
(A-43)
133
-------
Equations (A-38) and (A-41) with the substitutions indicated by
Equations (A-39), (A-40), (A-42), and (A-43) are solved to yield values
for A and C so that the equations for the magnetic potential finally
become
^ = (_HO + |)r Cos e (A-44)
and
Cos 6 " V Cos
To determine the force on the particle in the magnetic field, we need
the magnetic field outside the wire or
-*• ->
H2 = - V$2 (A-46)
where
or
o r + —. e + —=- z
2 9r r 96 8z
H2 = (^-y Cos 6 4 HQ Cos e)r - 1 (-|^ Sin 6 + HQr Sin e)e. (A-48)
Recall that
or
Fm = (Frr + FB9)- (A-49)
134
-------
Performing the appropriate operations with the results from Equations (A-44)
and (A-45) yields
(A-50)
and
Sin 29
(A-51)
The equation of motion for the particle in the air stream and in the
magnetic field can be obtained for the coordinate system established in
Figure A-l.
The air stream velocity has components'
-*•
V = V
o or
Voe6
(A-52)
where assuming potential flow,
vor = * V1 - ¥ Cos e
r
(A-53)
and
(A-54)
The Lagrangian form of the equation of motion in cylindrical coordinates
is
1" [•
m d2r - m r/de\2 | : + 1 m r d^e + „ _dr de
9 •
(A-55)
135
-------
where F are the aerodynamic forces and F are the magnetic forces.
The aerodynamic forces are assumed to be adequately expressed by Stokes
law
Fa = 6 TT n R V
a
(A-56)
where V is the relative velocity of the fluid and the particle.
Substituting for the aerodynamic and magnetic forces and rearranging
yields for the radial component (r)
,4 D3v
(I * R }Pp
6 ir n 1
! 1
V 9
I
Cos
a ,.<„.«._ ....._... _
2
dt
o Xy°(* 1
r/de ,
r^dt J
ir R3)M
6 y n Ra(f)3
a
-
+ a "dT~
M J- U P«^ OQ
o ' ' 'rt ^v»* *•• v
= V,
wherep is the particle density.
Defining r, = — and
a a
(A-57)
'm^ 2
a/ 9
X Vo H H0 R2
L ™2
(A-58)
Equation (A-57) can be rearranged to
v
o n
a u
2PpR2
9n
d ra . de ^2
dt2 al dt } _
dr.
V /
1 \ r m 1 / M
^ a rj l*W
+ Cos 2 e1
(A-59)
136
-------
Correspondingly, the angular component (e) is given by
204.
* Q i\
9 n
V
0 /
\r d2e
a dt2
i , 1
"a (l ' 2
dr
x o / S\ /d6 \
+ 2 (dt )(dt}
vm
} 9in fl m
. „ de
+ ra dt
Sin 29
; bin o a 3
(A-60)
Equations (A-59) and (A-60) constitute the equations of motion for
a paramagnetic particle under the influence of inertial, viscous, and
magnetic forces.
137
-------
TECHNICAL REPORT DATA
(Please read Inunctions on the reverse before completing)
. REPORT NO.
EPA-600/2-77-230
2.
3. RECIPIENT'S ACCESSION NO.
4. TITLE AND SUBTITLE
Application of High-Gradient Magnetic Separation
to Fine Particle Control
5. REPORT DATE
November 1977
6. PERFORMING ORGANIZATION CODE
. AUTHOR(S)
C. H. Gooding, T. W. Sigmon, and L. K. Monteith
8. PERFORMING ORGANIZATION REPORT NO.
9. PERFORMING ORGANIZATION NAME AND ADDRESS
Research Triangle Institute
P.O. Box 12194
Research Triangle Park, North Carolina 27709
10. PROGRAM ELEMENT NO.
1AB012: ROAP 21ADL-029
11. CONTRACT/GRANT NO.
68-02-1879
12. SPONSORING AGENCY NAME AND ADDRESS
EPA, Office of Research and Development
Industrial Environmental Research Laboratory
Research Triangle Park, NC 27711
13. TYPE OF REPORT AND PERIOD COVERED
Final: 6/75-8/77
14. SPONSORING AGENCY CODE
EPA/600/13
is. SUPPLEMENTARY NOTES lERL-RTP project officer for this report is Dennis C. Drehmel,
Mail Drop 61, 919/541-2925.
16. ABSTRACT
The report gives results of an assessment of the potential use of high-
gradient magnetic separation (HGMS) as a means of collecting gas stream particu-
lates. The assessment included both experiments and analyses of theoretical models.
Phase I included evaluations of theoretical explanations of HGMS and published
reports of liquid system applications. A bench-scale apparatus was constructed, and
HGMS experiments were conducted using redispersed dust from a basic oxygen fur-
nace.5 High efficiency collection of fine particulates was achieved with both high
throughput and reasonable projected energy rqeuirements relative to conventional
devices. In Phase n, experiments were scaled up to 0. 8 cu m/s (1700 CFM). Dusts '
from basic oxygen and electric arc furnaces were redispersed and collected. Results
show that submicron particles can be collected with 90-plus % efficiency using applied
magnetic flux densities of 0.2-0.4 T; With superficial gas velocities up to 11 m/s, the
pressure drop across the HGMS device was typically less than 1. 5 kPa (6 in. H2O).
Even lower fields can be used successfully at the expense of higher pressure drop or
reduced throughput.
17.
KEY WORDS AND DOCUMENT ANALYSIS
a.
DESCRIPTORS
b.lDENTIFIERS/OPEN ENDED TERMS
c. COS AT I Field/Group
Air Pollution
Dust
Magnetic Separators
Air Pollution Control
Stationary Sources
Particulate
High-gradient Magnetic
Separation
13B
11G
3. DISTRIBUTION STATEMENT
Unlimited
19. SECURITY CLASS (This Report)
Unclassified
21. NO. OF PAGES
20. SECURITY CLASS (Thispage)
Unclassified
149
22. PRICE
EPA Form 2220-1 (9-73)
138
-------