EPA-650/2-74-119-CI
OCTOBER 1974
Environmental Protection Technology Seri
es
i::^!:::^
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EPA-650/2-74-119a
ENTRAINMENT SEPARATORS
FOR SCRUBBERS -
INITIAL REPORT
by
S. Calvert, I.L. Jashnani, S. Yung,
and S. Stalberg
A.P.T. , Inc.
P.O. Box 71
Riverside, California 92502
Contract No. 68-02-0637
Program Element No. 1AB013
ROAPNo. 21ACX-086
Project Officer: L.E. Sparks
Control Systems Laboratory
National Environmental Research Center
Research Triangle Park, North Carolina 27711
Prepared for
OFFICE OF RESEARCH AND DEVELOPMENT
ENVIRONMENTAL PROTECTION AGENCY
WASHINGTON, D.C. 20460
October 1974
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This report has been reviewed by the Environmental Protection Agency
and approved for publication. Approval does not signify that the con-
tents necessarily reflect the views and policies of the Agency, nor does
mention of trade names or commercial products constitute endorsement
or recommendation for use.
11
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PREFACE
This report, "Entrainment Separators 'for Scrubbers",
is the interim report submitted to the Control System
Laboratory for E.P.A. Contract No. 68-02-0637.
The scope of work of this experimental and theoretical
study was to
1. Evaluate present technology.
2. Conduct an experimental study of air water
systems aimed at simulating the performance
of various types of entrainment separators.
3. Develop improved engineering equations and
methods for entrainment separator selection.
4. Develop and evaluate on a small pilot basis
new entrainment separator designs.
5. Develop specific research and development
recommendations.
Dr. Leslie E. Sparks, of the Control Systems Laboratory,
National Environmental Research Center, Environmental Pro-
tection Agency, was the Project Officer for this program.
Dr. Seymour Calvert, of A.P.T., Inc., was the Project
Director,
iii
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ACKNOWLEDGEMENT
A.P.T., Inc. wishes to express its appreciation for
excellent technical coordination and for very helpful
assistance in support of our technical effort to
Dr. Leslie Sparks, E.P.A. Project Officer.
IV
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TABLE OF CONTENTS
Page
Acknowledgement iv
List of Figures ix
List of Tables xvii
Nomenclature xix
Abstract xxv
Sections
Introduction 1-1
Summary and Conclusions 2-1
Literature Survey 3-1
Collection Mechanisms 3-3
Inertial Impaction 3-3
Sedimentation 3-6
Entrainment Drop Diameters 3-6
Design Equations for Entrainment Separators. • 3-9
Cyclone 3-10
Primary Efficiency 3-10
Pressure Drop 3-13
Packed Bed 3-14
Primary Efficiency 3-15
Pressure Drop 3-16
Zigzag Baffles 3-16
Primary Efficiency .... 3-16
Pressure Drop 3-21
Tube Bank 3-21
Primary Efficiency 3-21
Pressure Drop 3-23
Mesh 3-23
Primary Efficiency 3-24
Pressure Drop 3-25
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Page
Sieve Plates 3-27
Primary Efficiency 3-27
Pressure Drop 3-27
Reentrainment 3-31
Transition from Separated to Separator -
Entrained Flow 3-32
Horizontal Flow in Circular Tube • • • 3-33
Vertically Upward Flow in a
Circular Tube 3-36
Vertically Downward Flow in a
Circular Tube 3-38
Effect of Impingement of Gas Jet . . . 3-38
Interfacial Waves 3-38
Entrained Fraction and Rate of Entrainment 3-47
Drop Diameter of Reentrainment 3-50
Reentrainment Due to Rupture of Bubbles. . . 3-51
Drop Diameter of Reentrainment 3-52
Formation of Jet After Bubble Burst . . 3-55
Trajectory 3-57
Creeping of Fluids 3-60
Shattering of Drops 3-62
Manufacturers' Survey 3-62
Theory 4-1
Zigzag Baffles 4-1
Primary Efficiency 4-1
Pressure Drop 4-8
Reentrainment 4-8
Horizontal Baffles 4-11
1. Derivation of (X 4-11
2. Additional Equations 4-15
Vertical Baffles 4-16
1. Liquid Flow Due to Pressure Gradient4-18
2. Liquid Flow Due to Gravity 4-19
3. Additional Equations 4-20
VI
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Page
Cyclone 4_20
Primary Efficiency 4-20
Reentrainment 4-26
Onset of Reentrainment 4-26
Drop Diameter of Reentrainment 4-27
Rate of Reentrainment 4-27
Reentrainment Calculation 4-28
Auxiliary Experiment 5-1
Sampling for Droplet Size Distribution 5-2
Discussion 5-6
Results 5-8
Conclusions 5-17
Experimental Pilot Plant 6-1
Description of the Pilot Plant 6-1
Air Prefilter 6-2
Blower 6-2
Air Heater 6-2
Spray Section 6-2
Observation Sections 6-6
Drainage of Liquid in the Test Section • • • 6-9
Liquid Catch Tanks 6-10
Liquid Supply Tanks 6-10
Control Panel for Equipment 6-10
Electrical Supply Panel 6-11
Water Supply 6-11
Test Section 6-11
Mesh 6-12
Packed Bed 6-13
Zigzag Baffles 6-13
Cyclone . 6-13
Tube Bank 6-17
Calibration 6-17
Experimental Procedure 6-20
Sampling Procedure 6-21
Problems and Possible Errors in the Experiments • 6-28
vii
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Page
Experimental Results and Discussion 7-1
Inlet Entrainment 7-1
Experimental Results 7-9
Packed Bed 7-10
Overall Efficiency 7-10
Pressure Drop 7-10
Reentrainment 7-10
Zigzag Baffles 7-16
Overall Efficiency 7-16
Pressure Drop 7-21
Reentrainment 7-21
Liquid Flow on the Baffles 7-29
Mesh 7-33
Overall Efficiency 7-33
Pressure Drop 7-35
Reentrainment 7-35
Visual Observation of Reentrainment . . 7-42
Tube Bank 7-44
Overall Efficiency 7-44
Pressure Drop 7-49
Reentrainment 7-53
General Observations 7-S3
Pressure Drop 7-53
Reentrainment 7-66
Future Research and Development Recommendation . . . . 8-1
Appendix
Glossary
References
Vlll
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LIST OF FIGURES
Page
3-1 Entrainment Separator 3-2
3-2 Theoretical Impaction Efficiency as a
Function of Inertial Parameter for
Different Targets 3-2
3-3 Terminal Settling Velocity and Reynolds
Number for Water Drops in Air at 20°C
and 760 mm Hg 3-8
3-4 Cyclone with Tangential Gas Inlet 3-11
3-5 Generalized Flooding and Pressure Drop
Correlation for Packed Beds 3-17
3-6 Theoretical and Experimental Collection
Efficiencies of Rectangular Aerosol Jets . . . 3-22
3-7 Friction Factor, "f", Versus Reynolds
Number, "N ", for Wire Mesh Entrainment
Separator e Without Entrainment Load .... 3-26
3-8 Pressure Drop Due to Presence of Liquid in
the Knitted Mesh Wire with the Crimps in
the Same Direction 3-28
3-9 Pressure Drop Due to Presence of Liquid in
the Knitted Mesh with the Crimps in the
Alternate Direction 3-29
3-10 Baker Chart 3-34
3-11 Flow Pattern Diagram for Horizontal Flow • • . 3-35
3-12 Onset of Entrainment in Air-Water Flow .... 3-37
3-13 Comparison of Entrained Fraction in Upward
and Downward Co-current Annular Flow in
1.25 cm I.D. Tube 3-39
3-14 Extrapolation Method for Determination of
Point of Onset of Entrainment for Vertical
Downflow in 2.2 cm I.D. Tube 3-40
IX
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No.
3-15 Region of Wave Flow in Air-Water Flow in
a 3.2 cm Bore Tube 3-42
3-16 Relationship Between Effective Roughness
Height and Film Thickness 3.45
3-17 Breakdown of Disturbance Wave by Undercutting. 3-46
3-18 Breakdown of Disturbance Wave by Rolling . . . 3-46
3-19 Correlation of Entrainment Data by Hughmark. . 3-49
3-20 Histogram Showing Size Distribution of
Large and Small Drops Resulting from
Bubble Burst 3-54
3-21 Sauter Mean Diameter D Against Bubble
Diameter at 25°C . . .s 3.54
3-22 Time Against Velocity Curves for Drops
of Varying Degree 3-59
3-23 Height Against Velocity Curves for Drops
of Varying Diameter 3-59
3-24 Calculated Collection Efficiency for
Water Droplets in Air for ACS Mesh 3-64
3-25 Pressure Drop Versus Air Velocity for
10 cm Thick ACS Style 4CA Mesh 3-65
3-26 Pressure Drop Versus Steam Velocity at
Various Pressures through 10 cm Thick
ACS Mesh 3.66
3-27 Collection Efficiency Versus Particle
Diameter for Air Purification Method
Cyclone Separator with Inlet Velocity
of 16-21 m/sec 3-68
3-28 Comparison of Pressure Drop Characteristics
of Beco Engineering Type E/N-1 Pad and
Conventional Pad 3-70
3-29 Efficiency Curves for Burgess Industries
Centrifuge 3-72
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No. Page
3-30 Pressure Drop Characteristics of 10 cm
Thick Nu-Standard and 15 cm Thick Hi-
Thruput Mist Eliminators 3-73
3-31 Safe Operating Limits for Nu-Standard
and Hi-Thruput Mist Eliminators 3-74
3-32 Flooding Correlation for Fibrous Bed
Structure by Porter and Lucas 3-77
3-33 Performance of Koch Fleximesh Separator. . . . 3-80
3-34 Collection Efficiency of Particulate
Collectors . 3-81
3-35 Efficiency Comparison of York Scrubber
with Two Stage Wire Mesh Mist Eliminator . . .3-83
4-1 Continuous Zigzag Baffles 4-2
4-2 Drag Coefficient Versus Reynolds Number
After Foust et al. (1959), with Sphericity
IMF " as the Parameter 4-5
4-3 Penetration Versus Gas Velocity for
Baffle Section 4-6
4-4 Comparison of Primary Efficiency Curves
Based on Theoretical Model for Complete
Mixing and For No Mixing Models 4-7
4-5 Drag Coefficients for Flow Past Inclined
Flat Plates 4-9
4-6 Forces on An Element of Liquid Film on
Baffle 4-12
4-7 Effects of Baffle Edges on Reentrainment . . . 4-12
4-8 Maximum Liquid Load, QT/Qr, Versus Gas
Velocity in Horizontal b Baffle Section. . . 4-17
4-9 Predicted Effect of Liquid Loading in
Inlet on Reentrainment from Vertical Baffles . 4-21
4-10 Predicted Effect of Liquid Loading in
Inlet on Reentrainment from Vertical Baffles . 4-22
xi
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' Page
4-11 Cross Section of a Cyclone .......... 4-25
4-12 Predicted Effect of Inlet Gas Velocity
on Reentrainment in Cyclone with
a = 9.2 cm, b - 28 cm, d = 127 cm,
h= 132 cm, and d = 67 cm
4-29
4-13 Predicted Effect of Liquid Loading in
Inlet on Reentrainment in Cyclone with
a = 9.2 cm, b = 28 cm, d = 127 cm,
hs = 132 cm and dg = 67 cm .......... 4-30
4-14 Predicted Effect of Inlet Gas Velocity
on Efficiency of Cyclone with a = 9.2 cm
b = 28 cm, d = 127 cm, h = 132 cm and
de = 67 cm ' ...... • .......... 4-31
4-15 Predicted Collection Efficiency With and
Without Reentrainment for Cyclone with
a =9.2 cm, b = 28 cm, d = 127 cm,
hg = 381 cm, and d = 67ccm ......... 4-33
5-1 Apparatus to Study Onset of Entrainment
Velocities .................. 5.3
5-2 Predicted Effect of Drop Diameter on
Capture Efficiency by 5 cm Diameter
Filter Paper held Perpendicular to the
Air Velocity ................. 5.7
5-3 Distribution of Droplet Size in Air-Water
Entrained-Separated Flow ........... 5-9
5-4 Distribution of Droplet Size in Entrained-
Separated Flow on Log-Probability Graph. . . . 5-10
5-5 Comparison of Drop Size by Different
Investigators ................ S-ll
5-6 Effect of Impingement Angle on Drop Size . . . 5-13
5-7 Effect of Liquid Flow Rate on Impingement
Angle for Onset of Entrainment ........ 5-14
5-8 Comparison of Entrainment Onset Velocity
by Different Investigators .......... 5-15
xii
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No. Page
6-1 Block Diagram of Experimental Apparatus. . . . 6-3
6-2 Top View of the Entrainment Separator
Pilot Plant 6-4
6-3 Flow Diagram Showing Connection Between
Various Tanks and Pumps 6-5
6-4 Nozzle Position in the 30.5 cm x 61 cm Duct. . 6-8
6-5 Calculated Collection Efficiency for
Water Droplets in Air 6-14
6-6 Pressure Drop Versus Air Velocity for
10 cm Thick ACS Style 4CA Mesh 6-15
6-7 Baffle Section 6-16
6-8 Cyclone Assembly 6-18
6-9 Test Section with Bank of Tubes 6-19
6-10 Form for Recording Test Data 6-22
6-11 Calibration of Whatman No. 1 Filter Paper. . . 6-25
6-12 Liquid Load Sampling 6-26
6-13 Sampling Device Consisting of Impactor,
Heated Inlet Probe, Dry and Wet Bulb
Thermometer and Accessories 6-29
7-1 The Effect of Gas Velocity on the Mass
Median Drop Diameter 7-3
7-2 Drop Diameter Versus Volume Percentage
for Hollow Cone Nozzle Spraying Water at
10.2 atm Gauge Pressure 7-6
7-3 Drop Diameter Versus Volume Percentage
for Hollow Cone Nozzle Spraying Water at
6.8 atm Gauge Pressure 7-7
7-4 Drop Diameter Versus Volume Percentage
for Fulljet Nozzles Spraying Water at
2.7 atm Gauge Pressure 7-8
7-5 Experimental Collection Efficiency of a
Packed Bed as a Function of Gas Velocity
with Horizontal Flow 7-11
Xlll
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No. Page
7-6 Dry Pressure Drop in Packed Bed Versus
Gas Velocity 7-12
7-7 Wet Pressure Drop in Packed Bed Versus
Gas Velocity 7-13
7-8 Maximum Drop Diameter in the Entrainment
Versus Gas Velocity for 30 cm Packed Bed . . . 7-14
7-9 Mass Median Drop Diameter Versus Geometric
Standard Deviation Downstream of 30.5 cm
Packed Bed 7-15
7-10 Collection Efficiency Versus Gas Velocity
in the Zigzag Baffle Device with n = 6
and 9 = 30° 7-17
7-11 Collection Efficiency Versus Gas Velocity
in the Zigzag Baffle Device with n - 6
and 6 = 30°, d - 380 ym and a = 1.52. . . . 7-18
Pg g
7-12 Collection Efficiency Versus Gas Velocity
in Zigzag Baffle Device with n = 6,
6 = 30°, d = 1,225 ym and a = 1.75.
Theory Predicts 100% Efficiency 7-19
7-13 Experimental Collection Efficiency as a
Function of Gas Velocity in the Vertical
Baffles. Solid Line Represents Theory .... 7-20
7-14 Dry Pressure Drop in Baffles Versus Gas
Velocity 7-23
7-15 Wet Pressure Drop in Baffles Versus Gas
Velocity 7-24
7-16 Outlet Drop Diameter Versus Gas Velocity
for Zigzag Baffles with Inlet Mass Median
Diameter of 380 ym 7-25
7-17 Outlet Drop Diameter Versus Gas Velocity
for Zigzag Baffle with 1,225 ym Inlet
Mass Median Drop Diameter 7-26
7-18 Size Distribution of Drops Leaving Baffle
Entrainment Separators 7-27
7-19 Maximum Outlet Drop Diameter in the
Entrainment Versus Gas Velocity for
Zigzag Baffles 7-28
xiv
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No. Page
7-20 Drop Diameter Versus Geometric Standard
Deviation at Zigzag Baffle Outlet 7-30
7-21 Effect of Gas Velocity and Liquid Load
on Performance of Baffle Type Separator. . . . 7-31
7-22 Some Observed Phenomena in Entrainment
Separator 7-32
7-23 Experimental Collection Efficiency of Wire
Mesh as a Function of Gas Velocity 7-34
7-24 Pressure Drop in Wire Mesh Versus Gas
Velocity with Liquid Load as Parameter .... 7-36
7-25 Outlet Drop Diameter Versus Gas Velocity
for Mesh with 82 ym Inlet Drop Diameter . . . 7-37
7-26 Outlet Drop Diameter Versus Gas Velocity
for Mesh with 1,225 ym Inlet Drop Diameter . . 7-38
7-27 Maximum Outlet Drop Diameter in the
Entrainment Versus Gas Velocity for Mesh . . . 7-39
7-28 Drop Diameter Versus Geometric Standard
Deviation for Mesh 7-40
7-29 Effect of Gas Velocity and Liquid Load
on Performance of Mesh 7-41
7-30 Onset of Reentrainment Velocity Curves
for Mesh 7-43
7-31 Collection Efficiency Versus Gas Velocity
in Tube Bank with n = 6, d = 84 ym and
a = 1.32 Pf 7-45
g
7-32 Collection Efficiency Versus Gas Velocity
in Tube Bank with d = 380 ym and a = 1.52 . 7-46
Pg g
7-33 Collection Efficiency Versus Gas Velocity
in Tube Bank 7-47
7-34 Dry Pressure Drop in Tube Bank Versus
Gas Velocity 7-50
7-35 Wet Pressure Drop in Tube Bank Versus
Gas Velocity 7-51
xv
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No. Page
7-36 Friction Factor Versus Gas Velocity in the
Bank of Tubes 7-52
7-37 Outlet Drop Diameter Versus Gas Velocity
for Tube Bank with 380 ym Inlet Drop
Diameter 7-54
7-38 Outlet Drop Diameter Versus Gas Velocity
for Tube Bank with 1,230 ym Inlet Drop
Diameter 7-55
7-39 Size Distribution of Drops Leaving
Tube Bank Entrainment Separators 7-56
7-40 Outlet Mass Median Drop Diameter Versus
Geometric Standard Deviation for Tube Bank . . 7-57
7-41 Maximum Drop Diameter in the Entrainment
Versus Gas Velocity for Tube Bank 7-58
7-42 Experimental Results Showing the Effect of
Gas Velocity and Liquid Load on Performance
of Tube Bank 7-59
7-43 Pressure Drop in the Empty Section Versus
Gas Velocity 7-60
7-44 Dry Pressure Drop Versus Gas Velocity in
Different Separators Used in Pilot Plant . . . 7-62
7-45 Pressure Drop Through Entrainment Separator
As a Function of Superficial Gas Velocity . . 7-63
7-46 Comparison of Pressure Drop Data of
Houghton and Radford with Present Data .... 7-64
7-47 Minimum Outlet Drop Diameter Versus Gas
Velocity 7-70
xvi
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LIST OF TABLES
No. Page
3-1 Summary of Available Information on
Entrainment Separators 3-4
3-2 Summary of Conventional Entrainment
Separators 3-5
3-3 Packing Factors, "F", for Dumped Pieces 3-18
3-4 Packing Factors, "F", for Grids and
Stacked Pieces 3-20
3-5 Knitted Mesh Specifications . 3-30
3-6 Mesh Parameters 3-67
6-1 Nozzles Used in Spray Section 6-7
7-1 Nozzles Used in Spray Section 7-2
7-2 Drop Size Analysis 7-5
7-3 Comparison of Baffle Type Entrainment
Separators 7-21
7-4 Comparison of Tube Banks „ . 7-48
7-5 Comparison of Wire Mesh Demister 7-65
7-6 Effect of Liquid Load on Reentrainment in
Different Entrainment Separators 7-67
7-7 Observed Minimum Drop Size in the
Reentrainment 7-69
7-8 Reentrainment Observed by Kotov 7-71
A-l Overall Performance of Entrainment
Separator A-l
A-2 Drop Diameters of the Entrainment Entering
and Leaving the Test Section A-8
A-3 Liquid Material Balance A-16
A-4 Pressure Drop Data A-23
xvii
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XV111
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NOMENCLATURE
Latin
A = fractional area of the entrained drops projected
onto a plane
A = total flow area, cm2
A' = measured value of the fractional area of a plane
occupied by the drops
a = acceleration due to centrifugal force, cm/sec2
a1 = cross-sectional area of all the tubes in one row, cm2
a^ = height of entrance to cyclone, cm
a~ = specific area of mesh; surface area of wires per unit
volume of mesh pad, cm2/cm3
a, = constant of Equation (3-22)
a4 = defined after Equation (3-52), constant
a5 = defined in Equation (3-57), constant
a, = defined after Equation (3-59), constant
B = empirical constant for Equation (3-41)
b = spacing between two consecutive baffles in the
same row, cm
b' = perpendicular distance between two consecutive baffles
in the same row, cm
b., = width of entrance to cyclone, cm
b = diameter of bottom part of cone, cm
6
C = wave velocity
C = parameter reflecting the shape of the cyclone
C1 = Cunningham correction factor
C, = drag coefficient
c = number concentration, #/cm3
c = initial number concentration, #/cm3
c, , c9 = constants, defined by Equation (7-1)
-L Lt
XIX
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d = bubble diameter, cm
d = duct diameter, cm
d, = blot diameter, cm
d/ = measured blot diameter, cm
d = cyclone diameter, cm
d>
d = packing diameter, cm
cl
dg = diameter of exit pipe of cyclone, cm
d__ = equivalent (hydraulic) diameter of liquid film, cm
eq
d, » sieve plate hole diameter, cm
d. = diameter of jet, cm
d. - size parameter, cm
d = knitted mesh wire diameter, cm
d = mass median drop diameter, cm
dn^ = number of drops removed
d = drop diameter, cm
d_a = aerodynamic drop diameter, cm (g/cm3) V2
(d £/dz). = frictional pressure gradient for liquid
flowing alone in a tube, dyne/cm3
(d .p/dz)p = frictional pressure gradient for gas flowing
alone in a tube, dyne/cm3
d = inlet mass median drop diameter, cm
r o
d = Sauter mean diameter, cm
j
d = diameter of large drops, cm
s, n,
d = diameter of small drops, cm
E = fractional collection efficiency of an entrainment
separator
xx
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F = packing factor
F, = foam density, ratio of clear liquid height
to total foam height
F = correction factor = 1,1
f = friction factor
fD = drag coefficient
fg = friction factor in the absence of liquid phase
f^ = fraction of the perforated open area in the plate
G = mass flow rate of gas, Kg/m2sec
G^ = mass flow rate of liquid, Kg/m2sec
Gg = mass flow rate of gas, Kg/m2sec
g = acceleration of gravity, cm/sec2
H^ = fractional liquid hold-up in the bed
h = overall cyclone height, cm
h, = height of a jet, cm
h^ = dry-plate head loss
h = head over the weir
h = residual pressure drop
h = height of vertical cylinder of cyclone, cm
h = weir height
W
j = ratio of channel width to packing diameter
K = inertial parameter
K = Kutateladze number
K, = constant
L = mass flow rate of liquid, Kg/m2sec
Si = length of baffle, cm
xxi
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fl^ = natural height of vortex in cyclone, cm
&2 = length of mesh pad in the direction of flow, cm
&_ = unbroken length of jet, cm
&4 = broken length of jet, cm
N = number of blots in a given area if the blots do
not overlap
N = number of stages in the tube bank
N1 = number of blots in a given area
NR „ = drop Reynolds number
Nne G = gas Reynolds number
^Re L = liquid Reynolds number
n = number of rows of baffles or tubes
n = vortex exponent
n, = number of semicircular bends in packed bed
n. = collection efficiency for a given particle diameter
•^ in one stage of rectangular jet impingement
P = dimensionless parameter defined by Equation (5-15)
P = pressure, dyne/cm2
P = pressure outside bubble, dyne/cm
2
, = pressure due to gas velocity, dyne/cm2
Pt = penetration
AP = pressure drop, cm W.C.
APj = pressure drop in absence of liquid, cm W.C.
APT = pressure drop due to presence of liquid, cm W.C,
L*
APwet " APdry + iPL
XXI1
-------
Q = volumetric flow rate, cm3/sec
QL = liquid flow rate due to gravity, cm3/sec
R = entrainment parameter defined by Equation (3-41)
r = radius, cm
S = distance traveled by drop, cm
S, = height of exit pipe inside cyclone, cm
T = temperature, °K
t = time
t, = blot thickness, cm
u^ = cyclone inlet velocity, cm/sec
UG = superficial gas velocity, based on empty duct, cm/sec
u^ = actual gas velocity, cm/sec
UQ = mean gas velocity, cm/sec
UGC = reentrai-nment velocity, critical gas velocity for
reentrainment, cm/sec
ut = terminal deposition velocity, cm/sec
u = tangential velocity, cm/sec
v, = velocity of gas through hole in sieve plate, cm/sec
vi = actual gas velocity near the tubes
v = relative velocity, cm/sec
W = mass flow rate, g/sec
WLE = mass flow rate of entrained liquid, g/sec
w = baffle width
w^ = weir length, m
X = Martinelli parameter defined by Equation (3-40)
Y = distance from baffle to any point in the liquid film, cm
Z = bed length, cm
xxi 11
-------
Greek
a = volumetric flow ratio
a = initial surface disturbance, cm
a = amplitude of surface disturbance of a jet, cm
-------
ABSTRACT
Entrained drops of liquid must be separated from the
gas leaving scrubbers. This report gives results of an
evaluation of present technology, experimental studies of
entrainment separator characteristics, and theoretical
analysis. A pilot plant with 85 m3/min (3,000 CFM) maximum
capacity was built and experiments performed on zigzag
baffle, knitted mesh, tube bank, packed bed, and cyclone
devices. Horizontal gas flow was used in all cases, and
the cyclone axis was vertical.
Drop size distribution measurements on the inlet and
outlet enable the determination of collection efficiency
and reentrainment as related to drop size. Pressure drop
as a function of gas flow rate is also reported. Mathe-
matical models for primary collection efficiency are satis-
factory but useful reentrainment models are not yet available.
An auxiliary experiment was made to determine reentrain-
ment from liquid sheets under the influence of an air stream.
Work is still in progress on vertical gas flow systems,
solid deposition, mathematical modeling, and the development
of improved designs.
xxv
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XXVI
-------
T-T
high. Liquid drainage from the entrainment separator is a
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related problem. Poor "drainage will contribute to increased
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effects of reentfainment and flooding.
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' .Separators0als6 suffer from plugging and caking due to
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particle deposition and 'solute precipitation. Slugging; and
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caking increase pressure drop'and corrosion rate. The re-
the separator this w. , ,
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1001 efficiency
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1-2
-------
C. Determine and evaluate the adequacy of existing
theoretical models and design methods for pre-
dicting the performance of entrainment separators.
D. Review and evaluate performance of all major
types of entrainment separators currently avail-
able. Assess the advantages, disadvantages and
limitations of each type of equipment.
E. Identify specific operating and maintenance prob-
lems associated with entrainment separators.
Particular attention will be paid to the problems
encountered in S09 scrubbing systems under develop-
w
ment in E.P.A. programs.
Experimental Study
Conduct an experimental study of gas-water systems
aimed at simulating the performance of various types
of entrainment separators in the presence of soluble
and insoluble particulate matter. The experimental
study will investigate such variables as efficiency,
pressure drop, reentrainment, velocity, plugging and
related problems.
Selection and Design
Develop improved engineering equations and methods for
entrainment separator selection and design.
Pilot Test
Develop and evaluate on a small pilot basis new en-
trainment separator designs which offer improved
performance, e.g., efficiency, pressure drop, velocity,
freedom from reentrainment, and freedom from plugging
and related problems, when compared with available
entrainment separator designs.
Recommendations
Develop specific research and development recommenda-
tions for improving wet scrubber entrainment separators
1-3
-------
1-4
-------
SUMMARY AND CONCLUSIONS
This program of investigating wet scrubber entrain-
ment separators has been partially completed at this time.
The study indicates that presently available entrainment
separators suffer from various shortcomings such as over-
designs which necessitate large equipment size, lower
operating velocities due to flooding or reentrainment,
lack of reliable operating data on industrial separators,
and plugging.
The objectives of this study are to (1) Evaluate
present technology, (2) Conduct an experimental study of
air-water systems aimed at simulating the performance of
various types of entrainment separators, (3) Develop
improved engineering equations and methods for entrainment
separator selection, (4) Develop and evaluate on a small
pilot basis new entrainment separator design, and (5) De-
velop specific research and development recommendations.
EVALUATE PRESENT TECHNOLOGY
A literature search was carried out to evaluate the
technology on wet scrubber entrainment separators. Visits
were made to several libraries,and the literature available
in the A.P.T. library and identified by an APTIC computer
search was reviewed. Manufacturers of entrainment separators
were contacted by mail and asked for information. Visits
were made to E.P.A. and T.V.A. facilities to identify the
specific operating and maintenance problems associated with
entrainment separators.
The existing theoretical and empirical models to predict
the performance of the entrainment separators were evaluated.
The criteria for this evaluation were soundness of derivation
and closeness of comparison with actual performance.
2-1
-------
Experimental Study
A pilot plant to study wet scrubber entrainment sepa-
rators was built. It has a gas flow capacity of 85 m3/min
(3,000 CFM) and it consists of prefilter, blower, heater,
spray section, observation sections, test section, various
supply and catch tanks, and auxiliary equipment. Five
types of entrainment separators, namely, mesh, tube bank,
packed bed, cyclone,and baffles were studied. Up to now,
air-water systems with cross flow have been used in the
experiments. Observations include collection efficiency,
pressure drop, reentrainment, flooding, drainage, drop
size distribution and other variables.
Selectj^on _and Design
Attempts are made to develop mathematical models where
needed. Models for determining the following were developed
in the present study:
1. Primary collection efficiency in zigzag baffle
type entrainment separators.
2. Pressure drop in zigzag baffle type separators.
3. Primary collection based on either complete
turbulent mixing or no mixing.
4. Reentrainment in vertical zigzag baffles.
5. Reentrainment in horizontal zigzag baffles.
6. Reentrainment in a cyclone.
Auxiliary Experiment
Equipment was constructed and experiments were conducted
to determine (1) Onset of entrainment velocities in sepa-
rated flow, (2) Drop size distribution of entrainment,
(3) Effect of duct dimension, and (4) Effect of air jets
impinging on liquid film at various angles.
2-2
-------
CONCLUSIONS
The theoretical model developed to determine
primary collection efficiency for zigzag baffle
type separator agrees fairly well with experi-
mental results. Thus, the flow through the
zigzag baffles can be described as a series of
alternating bends with inertial mechanism
responsible for separation.
The pressure drop in zigzag baffles can be
determined from drag coefficients for inclined
plates held in the flow.
The overall collection efficiency in zigzag
baffles can be increased by staggering the
baffles. The pressure drop may be reduced by
keeping 1-3 cm distance between rows.
Models based on turbulent mixing reach 100%
efficiency as an asymtote with increasing gas
velocity. On the other hand, models based on
no mixing reach 1001 efficiency as a straight
line on efficiency-versus-gas-velocity curve.
The assumption of turbulent mixing gives better
agreement with actual performance of entrainment
separators .
It is possible to predict the drop collection
efficiency of entrainment separators at low gas
velocities (under industrially used conditions),
at least on pilot plant scale, from the available
literature.
At high gas velocities, reentrainment is a defi-
nite problem. Reentrainment may take place by
various mechanisms such as: a) Transition from
separated flow to separated-entrained flow,
2-3
-------
b) Rupture of bubbles, c) Creeping of liquid
on the entrainment separator surface, and
d) Shattering of liquid drops resulting from
splashing.
7. Transition from separated flow to separated-
entrained flow depends upon gas velocity, liquid
Reynolds number and liquid properties. The tran-
sition does not depend upon the duct dimensions.
The drop size distribution is independent of the
duct dimensions. The average drop diameter re-
sulting from this transition is about 250 urn.
The reentrainment velocity is considerably
reduced if jets of air stream strike the liquid
film at an angle. Therefore, sharp angles should
be reduced to avoid reentrainment.
8. The mechanism of reentrainment in zigzag baffles
is tearing of the liquid sheets caused by high gas
velocities and shattering of liquid drops. Reen-
trainment in cross flow baffles with horizontal
gas flow should be less than in baffles with verti-
cal gas flow. Zigzag baffles inclined at 30° from
gas flow direction should have less reentrainment
than baffles inclined at 45° from gas flow direc-
tion.
9. The reentrainment mechanisms in packed bed and mesh
pad are shattering of drops and rupture of bubbles.
Reentrainment resulting from small drops (less than
40 ym) due to rupture of bubbles is insignificant.
10. The mass median drop diameter due to reentrainment
was determined to vary between 80 ym and 750 ym.
Large drops (above 200 ym) are present due to
shattering of drops.
2-4
-------
11. Sampling of liquid drops and entrainment needs
careful consideration, Due to large drop size
in the reentrainment, a sedimentation effect is
present.
12. Drainage arrangements should be designed with
caution. An underdesign will lead to liquid
creeping in the separator and an overdesign may
result in reentrainment, from drainage, due to
gas flow.
13. The capital cost and the operating cost of en-
trainment separators in wet scrubbers represent
a significant factor and therefore entrainment
separators should be given more design attention.
14. The experience with wet scrubbers in power plants
indicates that performance of the entrainment
separator is crucial to the whole operation of
the scrubber.
2-5
-------
2-6
-------
LITERATURE SURVEY
The literature survey and background information
discussed in this chapter are examples of the selected
material on entrainment separators. For brevity, all
the literature in the field of separators is not men-
tioned. The following material summarizes what can be
distilled from a survey of the background information
and is a concept of what was known at the onset of this
study.
Figure 3-1 describes the entrainment separator. To
determine the performance of the separator, one needs to
know inlet loading (concentration and drop diameter),
primary collection efficiency, reentrainment, secondary
collection, liquid drainage and pressure drop. These
factors may be considered from a design point of view.
The inlet loading coupled with operating conditions will
determine reentrainment and drainage. Also, different
entrainment separators are limited to certain drop dia-
meters, below which their efficiency falls off sharply.
It is necessary to have engineering equations to determine
primary collection efficiency and pressure drop.
If the reentrainment rate and the size distribution of
the reentrained drops are known, secondary collection may
be determined. At the time of this study, no work had been
reported on the rate of reentrainment. All one can deter-
mine at this point is the range of gas velocities at which
onset of reentrainment takes place. The size distribution
of the reentrained drops may be guessed by assuming the most
probable mechanism of reentrainment for a given separator.
Knowledge of the pressure drop through a system is
important in calculating the energy loss incurred and in
selecting the proper pumps and other auxiliary equipment
to overcome that energy loss.
3-1
-------
Inlet Loading
Concentration —
and Drop Diameter
Primary
Collection
Efficiency
to
Reentrain-
ment and
Secondary
Collection
Outlet loading
•* Concentration
and Drop Diameter
Liquid Drainage
Figure 3-1. Entrainment separator
-------
Information which has been obtained from the litera-
ture survey is presented in Tables 3-1 and 3-2. A summary
of the literature search, dealing with primary collection
efficiency, pressure drop, reentrainment, entrainment
separator problems and details of separators now in use,
follows.
COLLECTION MECHANISMS
Knowledge of the basic mechanisms of drop collection
is fundamental to an understanding of entrainment separa-
tors. Some of these mechanisms are inertial impaction,
interception, sedimentation, diffusion and electrostatic
precipitation. Normal industrial operating conditions for
wet scrubbers are such that exiting drop diameters are
larger than 50 ym. The design and operating conditions of
separators thus favor inertial impaction and sedimentation
as the principal mechanisms of collection.
Iner t i al Imp ac t ion
This is the major drop collection mechanism in wet
scrubber entrainment separators. When fluid approaches
an obstacle the fluid streamlines spread around it. At
the same time inertial forces carry drops across the stream-
lines so that the drops hit and stick to the obstacle. It
is assumed that all drops colliding with the obstacle adhere
to it.
Two factors determine impaction collection efficiency.
The first is the velocity distribution of the gas flowing
by the collector, which varies with the Reynolds number of
the gas with respect to the collector. The second factor
is the drop trajectory, which depends on the mass of the
drop, its air resistance, the size and shape of the collec-
tor, and the rate of flow of the gas stream.
3-3
-------
Table 3-1. SUMMARY OF AVAILABLE INFORMATION ON ENTRAINMENT SEPARATORS
Cyclone
Mesh type
Separator
Tube
Bundle
Packed
Beds
Sieve
Plates
Baffles
Mechani-
cally
Aided
Impingement
Type
Typical Inlet Loading
Concen-
tration
<2£/m3
<2.7xlO-2
cm3/sec
cm2 of
liquid
"
*<'
— —
Drop
Diameter
>5ym
> y
>3ym
>20ym
>3ym
'•
Primary
Collection
Efficiency
/
/
/
/
/
X
X
/
Pressure
Drop
/
/
/
/
/
X
X
/
Approx.
Velocity
for Reen-
trainment
15-50m/sec
5-15m/sec
< 20m/sec
1-1 . Sm/sec
=3m/sec
•*• *^ Ani/cor*
•J • till/ o cC
30-45m/sec
Reen-
trainment
Drop
Diameter
70-400ym
100-500ym
100-500ym
small drops
^SOym
large drops
-0 . 1cm
100-500ym
Design information available in literature
X Design information not avialable in literature
-------
Table 3-2. SUMMARY OF CONVENTIONAL ENTRAINMENT SEPARATORS
Entrainment
Separator
(Construction
Material)
Cyclone
(Steel, S.S.)
Mesh
(Any material
that can be
drawn into
wire)
Fiber Bed
(Needs fibers
l-10ym)
Packed Bed
(Packing of
any mat'l)
Baffles
(Any mat '!)
^
Max imum
Capacity
m3/sec
141
65
Minimum
Drop
Size
5ym (2ym
for small
cyclones
Sum (1-
5pm for
series
combina-
tion)
0.3ym
3pm
lOym
Gas
Velocities
m/sec
Maximum
15-20
0.3-5
0.025-0.15
0.75-2.5
2-3.5
».
Separation
Efficiency
for Minimum
Drop
95%
951
99%
85%
85-95%
Pressure
Drop
cm W.C.
5-15
0.5-3
10-15
5-10%
of bed
length
2-2.5
Liquid
Load
Maximum
2.5 x
io-3
g/sec-
cm2
From
gener-
alized
corre-
lation
5% of
gas flow
rate by
weight
Comments
Surfaces may
erode easily
Basically 3
different
mesh
1. Crimped in
alternate
direction.
2. Nested
double
layer.
3 . Wound .
(Can plug
easily)
Can plug
easily
Better
drainage in
cross flow
Vertical
baffles
give better
drainage
than horiz.
04
I
in
-------
The equations of motion, the size and shape of the
collector, the mass of the drop and operating conditions
may be used to determine target collection efficiency.
Figure 3-2, from Golovin and Putnam (1962), gives theo-
retical impaction efficiency as a function of the inertial
parameter for different targets.
Sedimentation
The second collection mechanism important in entrain-
ment separators is sedimentation. Figure 3-3, from Fuchs
(1964), is a plot of drop terminal settling velocity versus
drop radius. Drop diameters encountered in wet scrubber
entrainment may vary from 50 to 500 ym, and the terminal
settling velocity for these drops will therefore range from
0.1 to 2.0 m/sec. The gas velocities used in entrainment
separators vary from 1.0 to 12.0 m/sec; however, except for
cyclone-type separators, which operate at very high veloci-
ties, most operate below 4.0 to 5.0 m/sec. Therefore, sedi-
mentation can be expected to affect the separation of drops.
ENTRAINMENT DROP DIAMETERS
To determine primary collection efficiency, it is
necessary to know the inlet entrainment drop diameter, which
depends on the entrainment source. The drop diameter may be
obtained by direct measurement or by the following method
if, as is often the case in wet scrubbers, entrainment is
cuased by pneumatic atomization. A correlation presented by
Nukiyama (1938, 1940) gives Sauter mean diameter as:
o.s „ *«
A - 58,600 / a \ ,q? PL /, nno ^
-------
U
U
W
1.0
0.8
0.6
0.4
0.2
Rtclangular half body
(ribbon with waltt)
Cylinder
- I Ribbon normal
"" I to
. flow
Ellipsoid of
revolution 101
thick
Ellipiold
" revolu-
vi>,20»
ick.
N'ACA 65A004 at
lero ancle o{
attack A\ thick
low-drag symmet-
rical airfoil
Joukowskl IS! thick sym-
metrical airfoil a: terq
angle of attack
0.1 1 10
INERTIAL PARAMETER, K =
100
"
Figure 3-2 - Theoretical impaction efficiency as
a function of inertial parameter
for different targets.
Golovin § Putnam (1962)
3-7
-------
1,000
T-T 100
U
<
W
ex,
o
tat
o
1000 3000
DROP RADIUS, ym
Figure 3-3 - Terminal settling velocity and
Reynolds number for water drops
in air at 20°C and 760 mm Hg.
After Fuchs (1964)
3-8
-------
where d = Sauter mean diameter, cm
v = relative velocity, cm/sec
a = surface tension, dyne/cm
p, = liquid density, g/cm3
UL = liquid viscosity, poise
QL/QG = liquid to gas volumetric ratio
In wet scrubbers the liquid load, QL/Qg» varies between
1.33x10-" m3/m3 and 4xlO'3 m3/m3 (1-30 gal/1,000 CFM). Gas
velocities are on the order of 60-100 m/sec (200-300 ft/sec).
The drop diameters obtained under these operating conditions
are between 50 urn and 500 ym.
If the liquid entrainment is formed in any other manner,
the entrainment drop diameter may be obtained as explained
later in this chapter in the section for the appropriate
reentrainment mechanism.
DESIGN EQUATIONS FOR ENTRAINMENT SEPARATORS
Our examination of collection mechanisms yielded the
basis upon which equations for primary collection effi-
ciency have been developed. Methods for determining en-
trainment drop diameter, a key variable in the calculation
of primary efficiency, were also reviewed.
Another important factor in the design of entrainment
separators is the pressure drop. A knowledge of the pres-
sure drop through a separator is useful for the selection
of auxiliary equipment such as blowers and for calculating
the separator's operating costs.
Six different kinds of entrainment separators will be
discussed in this section: cyclone, packed bed, zigzag
baffles, tube bank, mesh and sieve plate. A brief descrip-
tion of each is given, after which, equations for primary
efficiency and pressure drop are presented. These equations
3-9
-------
were selected as the best available expressions, based on
agreement with experimental data and soundness of deriva-
tion. Calvert et al. (1972) have evaluated some of these
equations.
The following assumptions are generally made in the
derivation of the primary collection efficiency equations:
1. A drop is spherical in shape.
2. The motion of a drop is not influenced by the
presence of neighboring drops.
3. All drops striking the collector surface adhere.
Cyclone
Commercially available cyclones in standard designs
for entrainment separators have a maximum capacity of up
to 141 m3/sec (300,000 CFM) of gas. Efficiencies of about
951 are claimed for 5 ym diameter drops in a well-designed
cyclone. Some manufacturers use a bundle of small cyclones
(multicyclones), which can efficiently collect drops as
small as 2 ym in diameter. However, this arrangement re-
duces the capacity of the device.
Primary Efficiency - Leith and Licht (1971) derived an
equation to predict primary collection efficiency in conical
bottom cyclones as pictured in Figure 3-4. With slight modi-
fication it can be applied to cylindrical cyclones. The
following assumptions were made:
1. The drag force in the radial direction on the drop
is given by Stokes law.
2. The tangential velocity component of the drop is
related to the radial position by a modified form
of the equation for a free vortex in an ideal fluid
ut rn = constant (3-2)
3-10
-------
Figure 3-4. - Cyclone with tangential gas inlet.
3-11
-------
3.
where "r" is the distance from the vertical axis
of the cyclone and "n" is defined below in
equation (3-5).
Backmixing of the drops takes place in the gas phase.
The primary efficiency for a conical bottom cyclone is
expressed as:
E = 1 - exp
f"
2 (CV
(3-3)
TT d
C =
2(1 -
1 +
h
Sl - V
- s.
(3-4)
0.3
(0.393 d )
2.5
(3-5)
S! * *i - hs
h - h
(3-6)
H, = 2.3 d ,
1 e \an
1/
1/3
(3-7)
= K
n+1
(n + 1)
(3-8)
3-12
-------
where E = fractional collection efficiency of an
entrainment separator
C = parameter reflecting the shape of the cyclone
H! = size or operating characteristic
n = vortex component
d = cyclone diameter, cm
a-j^ = height of the entrance to the cyclone, cm
b, - width of the entrance to the cyclone, cm
d = gas exit pipe diameter in cyclone, cm
S, = height of the exit pipe inside the cyclone, cm
£, = natural height of vortex in cyclone, cm
h = height of vertical cylinder of cyclone, cm
j
d. = size parameter, cm
T = temperature, °K
b = diameter of bottom part of cone, cm
e f
h = overall cyclone height, cm
K = inertial parameter
a = drop density, g/cm3
d = drop diameter, cm
u = tangential velocity, cm/sec
yfi = gas viscosity, poise
Pressure Drop - Because cyclones are operated at higher gas
velocities than other entrainment separators, the pressure
drop in a cyclone is usually higher than in other devices.
Cyclones are sometimes equipped with inlet vanes, which
serve to introduce the gas stream to the cyclone more
smoothly. The pressure drop in a device with an inlet
vane is expected to be lower than in one without. Shepherd
and Lapple (1940) derived an equation for a cyclone with
inlet vanes for pressure drop as a function of inlet gas
velocity and the cyclone inlet and outlet dimensions:
3-13
-------
AP = 0.00513 pG
where PG = gas density, g/cm3
Equation (3-9) can be modified by writing it as a
function of the geometric average of the gas velocity at
the cyclone inlet and outlet:
AP = 0.00513 x 5.89 p v (3-10)
Shepherd and Lapple also developed an equation for a
cyclone without inlet vanes:
/ Q \2/16 a,b \
AP = 0.00513 p
-------
Primary Efficiency - Jackson and Calvert (1966) and Calvert
(1968) have developed a theoretical relationship between
particle collection efficiency and packed bed operating
parameters by considering the bed to be made of a series
of semicircular channels. Their formulation included the
following assumptions:
1. The drag force on the drop is given by Stokes Law.
2. The number of semicircular bends, "n,", is related
to the overall height, "Z", of the packed section
of the bed, the packing diameter, "d ", and the
channel width, "b,", when any consistent units
may be used, by:
ni • 3-ri- CJ-12)
These assumptions led to the following equation for
predicting the primary efficiency for a packed bed
E = 1 ' ""Hd + j*' (. . Hd) ajY KP1 <3-13'
i * 7T- (3-14)
KP • r
where j = ratio of channel width to packing diameter
= bed porosity
Hd = fractional liquid hold-up in the bed
Z = bed length, cm
^cl = Packi-ng diameter, cm
UG = superficial gas velocity, cm/sec
dpa = aerodynamic drop diameter, cm(g/cm3)^2
3-15
-------
Pressure Drop - Perry (1963) gives a generalized pressure
drop and flooding correlation plot which appears as Fig-
ure 3-5, where a dimensional group of function ,
pGpLg
(centipoise)02, is plotted against a dimensionless group
of function i- (-—•) , where "G" and "L" refer to the gas
and liquid mass flow rates respectively. Values for the
packing factor, "F", for dumped pieces, stacked pieces and
grids are given in Tables 3-3 and 3-4. If "F" is not known,
— may be used instead.
e3
The operation of packed beds is limited by flooding.
The flooding lines for dumped pieces, grids and stacked
rings are shown in Figure 3-5. Pressure drop should be
obtained by using the largest gas and liquid streams.
Zigzag Baffles
Baffles can efficiently separate drops greater than
10 ym in diameter, while some of the better designed
devices can separate drop diameters of 5-8 ym. Common gas
velocities are 2.0-3.5 m/sec, and the pressure drop for a
6-pass separator is about 2-2.5 cm W.C,
The most common baffle shape is zigzag with 3 or 6
passes. These can be fabricated from a continuous wavy
plate or each pass is separated, in which case the separa-
tion distance is normally smaller than the width of the
baffles. Cross-flow baffles are claimed to have higher
drainage capacity than countercurrent flow baffles.
Primary Efficiency - A theoretical equation to predict
primary collection efficiency in zigzag baffles is not
available. However, Sarokin, Demidova and Kuzman (1968)
give the condition for minimum entrainment in the exit
stream:
3-16
-------
Ci,
•
U
ft,
0.5
0.2
0.1
0.05
0.02
0.01
0.005
0.002
0.001
0.01
FLOODING
LINES
Pressure cbrop
meters ot
/Pressure Drop,
[inches H20 per]
xft. of packing/
H20 per
meter of
packing
(dimensionless)
Figure 3-5 - Generalized flooding and pressure drop
correlation for packed beds (Perry, 1963)
3-17
-------
00
Raschig rings,
ceramic
.16 cm wall
.32 cm wall
.63 cm wall
.95 cm wall
Raschig rings,
carbon
.16 cm wall
.32 cm wall
.63 cm wall
.79 cm wall
Raschig rings,
metal
.08 cm wall
.16 cm wall
Lessing rings,
porcelain
.32 cm wall
.63 cm wall
Lessing rings,
metal
.08 cm wall
.16 cm wall
Table 3-3. PACKING FACTORS, "F", FOR DUMPED PIECES (m2/m3)
Nominal size of packing, cm
[0.64] [0.95] [1.27] [1.59] [1.9] .[2.5] [3.2] [3.8] [5] [8] [10]
5,250 3,280
510
430 210
121 98
5,250
1,340
920 525
430 210
118
2,300 1,280 980 560 510 380
1,340 950 720 450 360 272 187 105
(800)
(360)
(1,060)
(630)
(472) (387) (295) (200)
-------
Table 3-3. PACKING FACTORS, "F", FOR DUMPED PIECES (m2/ni3) (continued)
Nominal size of packing , cm
[0.64] [0.95] [1.27] [1.59] [1.9] [2.5] [3.2] [3.8] [5] fs] [10]
Partition rings 262 190
Pall rings, 318 171 105 82
plastic
Pall rings, 230 158 92 66
metal
Berl saddles 2,950 790 560 360 213 148
Intalox saddles, 2,380 1,080 660 475 322 171 131 72
ceramic
Intalox saddles, 108 69 52
plastic
Super-Intalox, 200 100
ceramic
Pellerettes 150
Parentheses denote a value of a/e3, rather than empirical F.
-------
Table 3-4. PACKING FACTORS,"?" FOR GRIDS AND STACKED PIECES
(m2/m3)
Nominal size of packing, cm
Wood grid
Metal grid
Grid tiles
2.5
20
8.2
3.8
11
5
8.2
8 10
5.9 4.9
13
118
14 15
Checker brick,
e=0.55 135
Raschig rings,
ceramic
.63 cm wall 95 16
.95 cm wall 36 12.8
Raschig rings, 21
metal
Partition rings,
diameter
7.6 cm length (1,200) (725)
10.2 cm length (705) (410)
15.2 cm length (375)
Partition rings,
square set
7.6 cm length (690) (460)
10.2 cm length (450) (275)
15.2 cm length . (260)
Parentheses denote a value of a/e3, rather than empirical F.
3-20
-------
0.35 < k = 10 yr( ] < 0.40 (3-16)
where k = Kutateladze number
This equation was confirmed with experimental data by
the authors.
Pressure Drop - No equation for the pressure drop in the
zigzag baffle type entrainment separator is available.
Mathematical models for predicting primary efficiency and
pressure drop in this device will be presented in the next
chapter.
Tube Bank
Although tube banks have not been used in entrainment
separators, equations predicting primary efficiency and
pressure drop have been developed and are presented below.
Primary Efficiency - Calvert and Lundgren (1970) found
that the collection efficiency for closely packed rods is
given by the equation for rectangular jet impaction. The
collection efficiency of each stage of impaction can be
found in Figure 3-6. Each row of tubes except the first
represents one stage of impaction. "B" is used as a para-
meter in Figure 3-6 and is defined by:
3 = 2 H/b (3-17)
where b = jet orifice width
H - distance between orifice and impingement plane
"K ", the inertia parameter, is defined with drop radius,
"r ", rather than diameter as in Figure 3-2.
Efficiency for the bank of tubes is given by:
E = 1 - (1 - n.)N (3-18)
3-21
-------
1.0
Chow
Exp. •
s Theory
0
Figure 3-6 - Theoretical and experimental
collection efficiencies of
rectangular aerosol jets.
3-22
-------
where n. = collection efficiency for a given particle
diameter in one stage of rectangular jet
impingement
N = number of stages in the tube bank
If the tubes are widely spaced, the target efficiency,
"n", can be calculated from Figure 3-2. In this case the
efficiency for the entire tube bank is:
B - i - (i - n ^-)n (3-19)
where a1 = cross-sectional area of all the tubes in one
row
A = total flow area
n = number of rows
Pressure Drop - a properly designed tube bank is claimed to
offer low pressure drop. Radford and Houghton (1940) found
a lower pressure drop in tests using a bank of tubes than
in using zigzag baffles or knitted mesh. Moyers (1960) ex-
pressed the pressure drop per row as:
AP = 4.95 x 10"spGuG2'19 (3-20)
Mesh
Knitted mesh of varying density and voidage is widely
used for entrainment separators. There are basically three
different kinds of mesh: (1) Layers with crimp in the
same direction - each layer is actually a nested double
layer. (2) Layers with crimp in alternate directions -
this results in an increase in voidage, reduced sheltering,
a decrease in pressure drop per unit length and an increase
in target efficiency per layer (3) Spirally wound layers -
the pressure drop is lower by about 2/3 than in layers with
crimp in the same direction, but the creeping of fluids, which
contributes to reentrainment, is expected to be higher.
3-23
-------
Standard mesh 10-15 cm thick having a density of about
0.15 g/cm3 is used to remove drops larger than 5 ym in dia-
meter. Gas velocities range from 0.3 to 5 m/sec and liquid
flow rate is limited by the drainage capacity of the mesh
to 2.5 x 10"3 g/sec cm2 of mesh. A lower density mesh made
of standard wires is used when 10-20% higher flow rates are
desired.
Often two mesh type separators in series are used to
remove drops in the 1-5 ym diameter range. The first mesh,
normally made of fine wires, coalesces the small drops, and
the second mesh, made of standard wires, removes them. The
first mesh is operated beyond the flooding velocity and the
second under flooding velocity. A major disadvantage with
this arrangement is a pressure drop which may reach 25 cm
W.C.
Some manufacturers use two or three stages of mesh, the
first being coarser and the final being finer, to remove
large and small drops successively.
A mesh type separator has the advantage that it can be
made to fit vessels of any shape. Any material which can be
drawn into the shape of a wire can be used for fabrication.
However, mesh separators are limited in application because
they plug easily. This can be avoided by upstream washing,
which will decrease removal efficiency and increase pressure
drop.
Primary Efficiency - Bradie and Dickson (1969) present the
following expression for primary efficiency in mesh separa-
tors :
E = 1 - exp (-| TT a2H2 n) (3-21)
where a^ = specific area of mesh; surface area of wires
per unit volume of mesh pad, cm2/cm3
&2 = length of mesh pad in the direction of flow, cm
r\ <= collection efficiency of cylinder wire
3-24
-------
The collection efficiency of cylindrical wire "n" can
be obtained from Figure 3-2. The factor of 2/3 in the ex-
ponential was introduced by Carpenter and Othmer (1955) to
correct for the fact that all the wires in the knitted mesh
are not perpendicular to the flow. That factor is the ratio
of the projected area of wires perpendicular to the flow to
the cross-sectional area of wires along the wire length.
The maximum allowable gas velocity in the knitted mesh
is calculated from the equation of Souders and Brown (1934):
/Pi - PG\V2
ur mav = 30.5 a^ ( Ji- £) (3-22)
G,max 3 \ pg /
where "a," varies with operating conditions and mesh design.
For most cases, a_ = 0.35.
Pressure Drop - The additional pressure drop due to the
liquid load may exceed the pressure drop in a dry knitted
mesh. York and Poppele (1963) have suggested that the
total pressure drop in the knitted mesh is the sum of the
pressure drop in the dry knitted mesh and the pressure drop
due to the presence of liquid:
AP = APdry + APL (3-23)
where AP, = pressure drop in absence of liquid, cm W.C.
APr = pressure drop due to presence of liquid,
cm W.C.
"AP, " can be obtained from an "f" (friction factor)
versus "NR p" (gas Reynolds number) correlation presented
in Figure 3-7. Carman (1937) developed the correlation to
determine the pressure drop in solid granular materials as
a dotted line on the plot. A similar curve is presented
for wire mesh by Stasangee (1948) and Shuring (1946). York
and Poppele obtained the data for knitted mesh both with
3-25
-------
oS
ii
5.0
2.0
.0
03 0.5
a.
0.2 h
0.1
10
r- 1
Satsangee data (1948)
and Shuring data (1946)
•Carman correlation for
solid granular materials
Crimp directioa
alternated
Crimps in
same direction
_L
100
N
Re,G
1,000
UG/a2
10,000
Figure 3-7 - Friction factor, f, versus Reynolds
number, NRg, for wire mesh entrain-
ment separator without entrainment
load.
3-26
-------
the crimps alternated and in the same direction. Their
curves are plotted in Figure 3-7 and should be used in
determining pressure drop. If the specific area, "a2",
is not specified, it can be determined from the mesh
porosity, "e", and the knitted mesh wire diameter, "dm":
4 ,-, _s j
m
(3-24)
Pressure drop data due to presence of liquid are not
available for all operating conditions or for mesh of dif-
ferent styles. Values of "AP/1 obtained by York and Poppele
are presented in Figures 3-8 and 3-9, with liquid velocity
as the parameter. The specifications of the knitted mesh
used in the two figures are shown in Table 3-5.
Sieve Plates
Primary Efficiency - Taheri and Calvert (1968) derived an
equation for sieve plate primary collection efficiency:
1 - exp (-40 F* K )
where 0.30 < F.^ < 0.65
P " 9
(3-25)
(3-26)
(3-27)
where
F, -
Vi =
d, =
foam density, ratio of clear liquid height to
total foam height
velocity of gas through hole, cm/sec
hole diameter, cm
Pressure Drop - Perry (1963) has suggested that the pressure
drop in sieve plates can be calculated according to:
AP = hw + how + hdp + hr
(3-28)
3-27
-------
10
1.0
0.1
0.01
i—i i
liquid
velocity
I i L I I I
3 456 7 8 9 10 11 12 13
UG[PG/(PL"PG}] > cm/sec
Figure 3-8 -
Pressure drop due to presence of
liquid in the knitted mesh with
the crimps in the same direction
3-28
-------
u
6
o
10,
z r i i i i I I i i E
i.o.
0.
0.01
I I I I I
I I
I
34 5 67 8 9 10 11 12 13
'
ucf PG/(PL'PG}] ' • cm/sec
Figure 3-9 -
Pressure drop due to presence of i
liquid in the knitted mesh with
the crimps in the alternate direction
3-29
-------
Table 3-5. KNITTED MESH SPECIFICATIONS
Direction of
crimps
e
a, cm2 /cm3
wire diameter,
d , cm
c*
Figure 3-8
same
0.977
3.61
0.028
Figure 3-9
alternate
0.99
1.51
0.028
3-30
-------
where h = weir height = 4-9 cm, assume 5 cm, (3-29)
if unknown 0
hQW = head over the weir • 0.143 FW ~ (3-30)
i Pr vv,
hdp = dry Plate head loss = T^ (3-31)
C Jj ^
h_ = residual pressure drop = 0.013 water (3-32)
r PL
-~ = 1.14 [0.4 (1.25 - fh) + (1 - fh)2] (3-33)
w
where FW = column wall curvature correction factor = 1.1
QL = liquid flow rate, here in m3/hr
w^ = weir length, m
f^ = fraction of the perforated open area in the
plate
REENTRAINMENT
As a result of the design equations presented in the
previous section, one is able to determine the primary col-
lection efficiency for five types of entrainment separators.
However, the overall separator efficiency is found to be
less than the primary efficiency because of reentrainment.
Increasing overall efficiency means reducing reentrainment,
the achievement of which requires a knowledge of the para-
meters important in determining the extent of reentrainment.
Thus, engineering equations describing this process are
vital to improved efficiency.
One mechanism of reentrainment results from high gas
velocity. In order to avoid this hazard, entrainment
separators are operated at lower gas velocities, resulting
in the large size and high cost of this equipment.
3-31
-------
Another source of difficulty is that the theory of two-
phase flow, which can serve as the basis for developing the
desired equations, is itself not well developed.
Due to the complexity of the subject and the necessity
of starting from the basics of two-phase flow, the following
material provides a lengthy but useful base from which
equations predicting reentrainment might be developed.
Reentrainment from an entrainment separator may take
place by any one or more of the following mechanisms:
1. Transition from separated to separated-entrained
flow caused by high gas velocity.
2. Rupture of bubbles at the gas liquid interface
and subsequent drop formation.
3. Creeping of the liquid along the solid surface
and movement into the gas exit in the entrainment
separator.
4. Shattering of liquid drops due to impaction.
The last three mechanisms of reentrainment depend upon the
design of entrainment separators. The first mechanism
represents the upper limit of the operation of entrainment
separators.
Transition from Separated to Separated-Entrained Flow
Reentrainment may occur at high gas velocities due to
transition from separated to separated-entrained flow. In
simple geometries such as straight tubes, the transition
takes place at much higher velocities than those at which
entrainment separators are operated. Yet reentrainment is
observed in separators at the lower velocity. This is
caused by such phenomena as the impingement of the gas
stream onto the liquid at an angle and the presence of gas
jets. Also the flow pattern in the entrainment separator
is not so uniform as in circular tubes. However, due to
3-32
-------
the limited information available, we must consider two-
phase flow in a simple geometry such as the circular tube.
In the operation of entrainment separators, flows may
be horizontal, vertically upward or downward, or inclined.
The onset of reentrainment depends upon the flow direction,
flow geometry and the fluid properties. We will examine
reentrainment in the simple geometries for which literature
is available.
Horizontal Flow in Circular Tube - The best known and most
widely used flow regime map for horizontal flow is that of
Baker (1954), which is shown in Figure 3-10. The Baker
chart as modified by Scott (1963) is shown in Figure 3-11.
where "GL" and "G " stand for the mass flow rates of liquid
and gas, respectively, in kg/m2 sec. The dimensionless
parameters "X" and " V," represent allowances for the fluid
physical properties in the system and are defined by the
following:
v - (°JL\
1 \aL/
(3-34)
0.5
(3-35)
where p. = density of air, g/cm3
Pw = density of water, g/cm3
aw = surface tension of water, dyne/cm
CL = surface tension of liquid, dyne/cm
yw = viscosity of water, poise
The onset of reentrainment may be considered as a
transition from separated flow to entrained flow. The
curve separating the two regions is shown in Figure 3-10
3-33
-------
1,000
u
-------
1,000 LI i Mi i I III I I III I I HI I I '+
DISPERSED
ANNULAR
BUBBLY
OR
FROTH
STRATIFIED
i u> i i m T ii
0.1
0.1 1.0
10
10'
Figure 3-11 - Flow pattern diagram for
horizontal flow (After Baker,
1954, as modified by Scott,
1963).
3-35
-------
From that figure, it can be calculated that for an air-water
system, entrained flow starts at GG = 125 Kg/m2 sec when
GL/GG = 1 Kg/Kg. If GL/GG = 100 Kg/Kg, entrained flow
starts at GG = 37.5 Kg/m2 sec. This corresponds to a re-
entrainment velocity of 100 m/sec at a liquid loading of
1.33 x 10"3 m3/m3. The transition velocity drops to 30 m/sec
at a liquid loading of 0.133 m3/m3.
Vertically Upward Flow in a Circular Tube - Cousins et al.
(1965) studied the transition from separated to separated-
entrained flow in a 0.95 cm bore vertical tube. The flow
was in the upward direction. Their general conclusion is
that for thin liquid films at low liquid Reynolds numbers,
the gas velocity for entrainment increases rapidly with de-
creasing liquid rate, and a limiting liquid flow rate may
be reached,below which no entrainment occurs irrespective
of further increases of gas rate. At higher liquid flow
rates, both gas and liquid flow rates are important in
governing the onset of entrainment. Figure 3-12 from the
work of Cousins et al. (1965) illustrates this conclusion.
For the asymptotic case of a limiting liquid flow rate,
the onset of entrainment corresponds closely to the formation
of large disturbance waves in the liquid film. For lower gas
flow rates the existence of large disturbance waves is not a
sufficient condition for the onset of entrainment. Thus,
there is a region in which the gas velocity is sufficient to
give rise to large waves on the surface but is insufficient
to break these waves into droplets. Zhivaikin (1962) sug-
gested that the onset of droplet entrainment in upward flow
can be correlated by the equation:
.5
(3-36)
/ V
\>w)
3-36
-------
800
600
Hj400
-------
where u~ = critical gas velocity for entrainment
NR L = liquid Reynolds number
Vertically Downward Flow in Circular Tube - For downward flow
the onset of entrainment equations are presented by Chien and
Ibele (1962), Steen and Wallis (1964) and Zhivaikin (1962)
and appear in Chapter 4. Each investigator defined onset of
entrainment differently and thus their predictions for onset
of entrainment differ significantly. Their data were ob-
tained in a 2.5 cm bore tube.
Measurements of the amount of liquid entrainment in
upward and downward co-current two-phase annular flow are
presented in Figure 3-13. It can be seen that for upward
flow the quantity entrained passes through a minimum with
increasing air flow rate, while for downward flow entrainment
increases with total liquid flow rate.
Effect of Impingement of Gas Jets - As mentioned earlier, the
gas and liquid phases do not flow parallel in the entrainment
separator. Jets of gas are present, which may impinge on the
liquid film at various angles. The presence of gas jets,
their impingement on the liquid film at various angles, etc.
depend upon the entrainment separator design.
Wallis (1962) studied entrainment in ducts with various
inlet designs. The reentrainment velocity varies with inlet
design, from 1,800 cm/sec to 2,400 cm/sec. The data are
shown in Figure 3-14.
Interfacial Waves - The study of interfacial wave behavior is
important in the determination of transition from separated
flow to separated-entrained flow. Experimental and theoreti-
cal studies of wave behavior and its influence on other
phenomena are still at a very early stage of development.
The most advanced theoretical studies have been concerned
3-38
-------
40
H
2
w
30
2 20
10
Water Flow
Rate, cc/min
I
0 10 20 30
AIR VELOCITY, m/sec
Figure 3-13 - Comparison of entrained fraction in
upward and downward cocorrent annular
flow in 1.25 cm I.D. tube. Solid line
represents upflow and dotted line
represents downflow.
3-39
-------
H
S5
W
W
60
50
40
30
20
10
20
40
60
AIR VELOCITY, m/sec
Figure 3-14 - Extrapolation method for
determination of point of
onset of entrainment for
vertical downflow in 2.2 cm
I.D. tube. Water flow rate =
1,280 cm3/min. The curves are
shown for different inlet
designs. After Wallis (1962)
3-40
-------
with the problem of the initial formation of waves, rather
than their development and influence. However, the in-
stability of the waves represents the physical phenomenon
responsible for transition from separated to separated-
entrained flow. Thus, to understand the physical phenomena
responsible for reentrainment in entrainment separators, one
should look at the interfacial waves, breaking of the waves,
drag friction on the film due to gas flow, etc.
A flow map showing different regions of interfacial
structure observed in air-water flow at atmospheric pres-
sure is illustrated in Figure 3-15 (Hall-Taylor et al. 1963).
As the air velocity is increased, ripple waves and distur-
bance waves take place. These waves represent instability
of the interface. One theory that explains interfacial in-
stability is the Kelvin-Helmholtz theory.
Consider the case of the flow of two inviscid fluids
of different densities (say a gas and a liquid), separated
by a vertical interface on which a regular train of waves
is moving with a velocity, "C". As the gas flows around
the curves of the streamlines, centrifugal forces are set
up and these must be balanced by a pressure gradient in the
direction normal to the streamline:
- If- >G? CSG - C)2 (3-37)
where r = radius of curvature, cm
Ug = mean gas velocity, cm/sec
n = direction normal to the streamline
The pressure at the interface can be found by inte-
grating this equation between the interface and infinity.
It follows from Equation (3-37) that for surfaces of
positive curvature the pressure at the interface will be
less than the undisturbed pressure at infinity, whereas for
3-41
-------
o
-------
surfaces of negative curvature the pressure at the interface
will be greater than the value at infinity. Thus, the gas
flowing past the wavy interface generates an increased pres-
sure over the troughs and suction over the crests. A similar
line of argument obviously applies on the liquid side of the
interface, only the liquid exerts an outward pressure at the
crests and suction at the troughs.
For vertical flow the only forces opposing these normal
stresses are those due to surface tension. For a stable
interface condition, the surface tension stresses exactly
balance the effects of the normal stress. On the other hand,
the wave will grow in amplitude when the sum of the local
liquid and gas normal stresses exceeds the surface tension
stress. It can be further deduced that the thinner the
liquid film the greater the gas velocity needed to cause
an increase in amplitude of a given size wave.
It may be expected that the effect of waves will be to
increase the friction factor. To determine the liquid flow
rate, it is necessary to have interfacial friction factor.
Roberts and Hartley (1961) found, on plotting friction
factor as a function of liquid film thickness for a given
gas velocity, that the friction factor did not begin to in-
crease with film thickness until a certain value had been
exceeded. They were able to correlate the difference between
the interfacial friction factor and that for the same gas
flow rate in the absence of the liquid phase, by the equation:
fi " fG
eq
5 /2\'-5l
" *^ l£GJ J
6 a " » i (3-38)
3-43
-------
and, £. = T. T PG UG (3-39)
where f.^ = interfacial friction factor
fG - friction factor in the absence of liquid
phase i
6 = liquid film thickness, cm
d = equivalent (hydraulic) diameter, cm
T^ = interfacial shear stress, gm/cm sec2
Thus, for very thin liquid films there would be no sig-
nificant waves on the interface and no effective roughness.
For thicker films there would be a minimum instantaneous film
thickness corresponding to the troughs of the waves on the
surface.
Gill et al. (1963) calculated the equivalent roughness
height and plotted it as a function of the film thickness
defined by Roberts and Hartley. The correlation is shown
in Figure 3-16.
Hanratty and Woodmansee (1965) suggested that the onset
of reentrainment may result from suction on the wave by the
Kelvin-Helmholtz mechanism. This suction will lead to a
tearing of the wave tips.
A number of possible mechanisms have been suggested by
which transfer of droplets can be effected by the waves, but
at present there is no definite evidence to favor any parti-
cular one. Lane (1957) described the mechanism illustrated
in Figure 3-17. The gas starts to "undercut" the wave and a
round, open ended bubble begins to form. The bubble grows,
leaving a thick-ringed filament around its base and event-
ually breaks up into droplets. Once the breakup occurs, the
excess (dynamic) pressure inside the bubble gives rise to a
rapid radial transport of the droplets.
3-44
-------
U
0.01 0.02 0.03
FILM THICKNESS, cm
Figure 3-16 -Relationship between
effective roughness
height and film
thickness.
3-45
-------
Gas
•Time
Figure 3-17 -
Breakdown of Disturbance Wave by Undercutting
Time-
Figure 3-18 -
Breakdown of Disturbance Wave by Rolling
3-46
-------
An alternative form of breakup is illustrated in
Figure 3-18. A large amplitude wave on a shallow liquid
layer tends to steepen at the front and then to form a
breaking wave. If the gas velocity is very high, it might
be expected that the tips of the waves would be drawn out
into thin liquid sheets with subsequent breakup.
Entrained Fraction and Rate of Reentrainment - If the gas
velocity is increased beyond the onset of reentrainment,
some of the liquid that is collected on the entrainment
separator elements will be reentrained* To determine the
overall efficiency of a separator, one must know the
equilibrium entrainment fraction, rate of reentrainment,
and reentrainment drop diameter. . In this section we will
deal with entrained fraction and the rate of reentrainment.
There are no reliable methods at present for estimating
the fraction of liquid flowing as entrained droplets out of
an entrainment separator. Geometrical factors, inlet con-
ditions, physical and chemical properties and change of
conditions along the entrainment separator can all affect
the results obtained. Therefore, any correlations that have
been developed for entrained fraction estimation should be
treated critically.
Wickes and Dukler (1960) have attempted to correlate
the data for entrained droplet flow rate by means of the
Martinelli (1949) parameter "X":
X =
(dp./dz)
1/2
where (dpr/dz)L = frictional pressure gradient for liquid
flowing alone in a tube, dyne/cm3
(dp^/dz),, = frictional pressure gradient for gas
flowing alone in a tube, dyne/cm3
3-47
-------
A dimensional entrainment parameter "R" was obtained by
intuitive reasoning and was defined as:
3.8x103 B W u./ur
R • B G
where W.£ = the flow rate of entrained liquid, g/sec
B = empirical constant, 22 for smooth injection
and 13 for shock injection
Collier § Hewitt's (1961) data correlate fairly closely
about a line given by the following equation
X = 0.069 R0-39 (3-42)
Hughmark (1973) correlated the entrainment data of
Collier $ Hewitt, Alia, et al. (1965) and Wicks and Dukler.
The data are plotted in Figure 3-19. Empirical correlating
equations are:
yj < 36 a = 0 (3-43)
36 < y* < 42, a = -0.000442 + 0.000013 y* (3-44)
42 < y* < 60, a = -0.000625 + 0.0000172 y* (3-45)
60 < y* . a • 5xlO-8 (y*)22 (3-46)
O \3
WL PG
where a = rr= = volumetric flow ratio
WG PL
and W,, W_ are liquid and gas flow rates, g/sec
Anderson, Bellinger § Lamb (1964) obtained the rate of
interchange between entrainment and separated flow in horizon-
tal pipes. Their results show that the rate of interchange
(% of the total liquid flow per 30.5 cm length of pipe) can
be correlated as follows:
3-48
-------
10'
I02
l0"3
I04
I05
• COLLIER AND HEWITT (1961)
OALIA CT AL (1965)
oCCUSINS ET AL (1965)
10
100
1000
Figure 3-19 - Correlation of entrainment data
by Hughmark (1973).
3-49
-------
% interchange = 4 for NRe L > 2,570, NRe fi > 3,000 (3-47)
% interchange =0.5 for NRe L = 1,150, NRe G > 3,000 (3-48)
% interchange = 0 for NRg G < 3,000 (3-49)
Drop Diameter of Reentrainment - Another important factor,
besides the amount of reentrainment, is the drop diameter of
reentrainment. If a secondary entrainment separator is used
to remove the reentrained drops, it is necessary to know the
drop diameter of reentrainment to design the secondary
separator.
Few investigators have measured the drop size distri-
bution arising from separated flow. Wicks and Dukler (1966)
obtained data for co-current downward flow in a 15 cm x
1.8 cm rectangular vertical channel. The mass median drop
diameter varied between 150 and 450 ym. The smallest drop
diameter found in the entrainment was 45 ym.
Cousins and Hewitt (1968) obtained data during experi-
ments on the mass transfer of liquid droplets subsequent to
removal of liquid film. Photographs of the droplets were
obtained for a zone immediately adjacent to the channel wall.
The mass median drop size was between 90 ym to 175 ym. They
also found a minimum droplet size of 45 ym. Both sets of
data show a decrease in the mass median diameter with in-
creasing gas velocity.
Fraser et al. (1963) studied drop formation from liquid
sheets by an air blast at 90°. They observed that the
liquid sheet does not break down upon immediate impact with
the air stream but is deflected away from it. Waves are
initiated at the point of impact and the sheet breaks down
into drops through the formation of unstable ligaments. The
air velocity was varied between 103cm/sec and 6xl03cm/sec
and the liquid flow rate between 12.5 and 125 g/sec. The
3-50
-------
liquid drop size varied from 30 urn to 250 ym. They ob-
served that the average drop size increases with increase
in liquid flow rate and decreases with increase in air
velocity.
Reentrainment Due to Rupture of Bubbles
The second mechanism which leads to reentrainment is
rupture of bubbles. Again the important questions are at
what velocity reentrainment begins, what is the drop dia-
meter of the reentrainment, what is the amount of liquid
in the reentrainment, etc. The effect of operating
variables such as gas velocity, liquid flow rate, etc.
should also be determined.
The following section deals with the mechanism of
bubble rupture, the drop size distribution created by
rupture of bubbles, and the trajectory of these re-
entrained drops.
Rupture of bubbles is the main cause of the re-
entrainment of liquid drops into the gas phase in devices
such as sieve plate, bubble cap plate, packed bed, and
mesh type separators. The collapse of a bubble when
exiting from the liquid phase is associated with thinning
of the liquid film starting at the top part of the bubble.
The upper surface thins to the extent of becoming weak
enough to rupture. Rupture of the upper part of the bubble
film takes place when the film thickness if of the order of
0.1 ym, provided there are no external disturbance forces
leading to the rupture of films (Kitchener, 1964; Jashnani,
1971). The collapse of the bubble at the interface leads
to the release of surface tension energy which is converted
into kinetic energy. The kinetic energy is sufficient to
impart high velocities to liquid drops formed during this
process.
3-51
-------
Drop Diameter of Reentrainment - The liquid drops which are
formed during the rupture of bubbles are of two distince
diameter'ranges. Small drops (about 30 to 40 ym) are formed
from the top of the bubble surface, and large drops (about
1 mm) are formed due to the disturbances in the liquid phase
following the bubble rupture. Formation of the small drops
was observed by Stumpner (1936), and he attributes drop
formation to the rupture of the projecting dome of the
bubble into fragments.
Davis (1940) cites the work of Edgerton, Germeshausen
and Grier (1936) on the rupture of a soap bubble to estab-
lish that during collapse, the dome opens and forms an
initial perforation near the apex without creating drops.
He suggests that large drops are subsequently formed by the
breakup of the liquid jet which forms when the surrounding
liquid moves inward to fill the depression caused by the
bubble. The depth of the depression is a function of the
bubble size. Drops are more likely to form from small,
rather than large, bubbles. The work of Dombrowski and
Frazer (1954) shows that a thin liquid film on the top of
the bubble breaks up by the initial formation of a number
of perforations which subsequently expand to give a lace-
like structure. The liquid ligaments so formed are unstable
and break into small drops of varying size.
Drop formation due to bubble burst occurs in three
steps. The first step, the lifetime of the bubble at the
interface, lasts on the order of l/100th sec.or longer; the
actual bubble burst, the second step, takes a few micro-
seconds; and events subsequent to the bubble burst extend
over a few milliseconds.
Newitt et al. (1954) measured the drop size distribution
due to the rupture of a bubble, and a typical distribution is
3-52
-------
shown in Figure 3-20. They also observed that in all cases the
Sauter mean diameter of large drops,"d ", decreases nearly
S ,Jt
linearly and the number of large drops, "No"» increases with
decreasing bubble diameter. Furthermore, both "d 0" and
s, x>
"N " decrease with rising temperature, and the change in the
At
number of drops produced by the bursting of large bubbles
decreases markedly as the temperature is raised. Figure 3-21
shows the variation of the Sauter mean diameter "d *" with
s, x,
the bubble diameter at 25°C, using distance from the inter-
face as parameter.
Large drops are formed from jets of liquids which are
formed following bubble rupture. As the liquid surrounding
the bubble moves inward to fill the depression caused by the
bubble, it usually forms a jet of liquid. As the bubble
diameter is reduced, the internal pressure rises and the
energy available for jet formation increases. This gives
rise to jets of higher initial velocity and hence greater
unbroken length. Stuhlman (1932), experimenting with bubbles
of 0.25 cm diameter or less, showed that the resulting jets
became progressively thinner with decreasing bubble diameter
and break down into two or more drops per jet.
Garner, et al. (1954) found that 95% of the droplets
entrained in the evaporator were below 20 ym, but because of
their low mass they formed only a small fraction (about 2%)
of the total weight of the entrained liquid.
Droplets from the rupture of bubbles larger than 0.5 cm
diameter were almost entirely produced from the bubble dome.
The entrainment resulting from bursting bubbles increased by
20 fold when the bubble diameters were reduced from 0.6 cm
to 0.2 cm. Mass median drop diameters of 440 ym, 660 ym
and 820 ym were obtained from bubbles having 0.2 cm, 0.3 cm
and 0.4 cm diameter, respectively.
3-53
-------
w
s 30
C/3
w
25
20
15
G 10
w
& 5
w
pi
* 0
I
&
irar
±31
1
I
TTT
m
II
Sfi
•tr1
Jil!
mm
il
0
Size range of small
particles .micron
140 1 - 16
Size range of large
particles, cm
Figure 3-20 - Histogram showing size distribution of
large and small drops resulting from
bubble burst
.101
••64 cm
1.91 cm
3.17
cm
-->i4.44 cm
5.71 cm
30 35 40 45 50 55
BUBBLE DIAMETER, cm
Figure 3-21 - Sauter mean diameter, D p, against
bubble diameter at 25°C '
3-54
-------
Small drops travel outward from the point of burst at
angles of up to 80° from the vertical. Thus, some of the
drops will be collected on the walls.
Formation of Jet After Bubble Burst - To determine re-
entrainment in the gas phase, it is necessary to study
the kinetics of the liquid jet formed after the bubble
burst. The dimensions of the jet are important, as the
instability of jets produces drops which are reentrained
in the gas phase. The velocity of rise of the liquid jet
will give the large drops their initial velocity.
Davis (1940) gives the following equation for jet
velocity based on the pressure pulse following bubble
rupture:
u = 6 _£* (3-50)
°Ld
where t = time during which the impulse force is applied
(about 3 x 10'5 sec)
d = bubble diameter, cm
Equation (3-50) predicts velocities which are very small,
about 0.1 cm/sec. According to Newitt et al . (1954), when
the rupture of a bubble takes place, a partial vacuum is
created due to the impulse force, and therefore the absolute
pressure rather than pressure excess should be used as an
impulse force:
u = (3-51)
2 aLd
where P = pressure outside the bubble, dyne/cm2
Equation (3-51) assumes no correction for the change in
bubble diameter due to the contact with air by part of the
bubble surface.
3-55
-------
Newitt et al. give the following equation for the height
of the jet:
(3-52)
For air water system, a4 = 1 cm2, a = 73 dynes/cm,
g - 980 cm/sec2 and PL = 1 g/cm3 give h^ 0.404 cm. The
experimental jet height as obtained from photographs is
0.44 cm.
Weber (1931) has calculated the relationship between
the unbroken length of the jet and its velocity. In the
range of velocities created by 0.2 to 0.5 cm bubble diameters,
the relationship between velocity and length is linear.
Haenlein's (1931) experimental values confirm this. The
distance, "A j", of the unbroken jet is given by the expression:
*3 = t' u
where t1 = rise ___-._—_ £=!*$ — - - (3-53)
time of the unbroken _______
length of jet, sec ^-r-r-JI^—l-^-Ir-jr^r
The velocity obtained from Equation (3-51) should be
used in Equation (3-53) to calculate the length of the
unbroken jet.
If the length of the liquid jet is longer than the peri
meter of its cross-section, it becomes unstable and breaks
up. The breakup occurs at different lengths along the jet,
and it is this broken part which forms the liquid drop.
For non-viscous liquids Weber (1931) gives the most
favorable wave length for breakup of a jet as:
djet = 4'42 djet
3-56
-------
where d. = diameter of the jet, cm
jet
Equating a cylindrical length of jet, equal to 4.42 d. ,
with the volume of a drop formed, results in:
dp = 1.89 d.et (3-55)
Castleman (1931) gives the following expression, based on
the work of Rayleigh (1879, 1889) and Weber (1931), for the
time of break-up of a liquid jet:
t = iil log ^i (3-56)
a5 ao
(v 0-5
4d^—] (3-57)
where, d = bubble diameter, cm
a.^ = amplitude of surface disturbance of the
jet - 0.5 d.Qt
a = initial surface disturbance = 10~5 cm
The above equation gives times which are roughly 4 times
greater than actually observed.
The total length of jet will be the sum of unbroken and
broken parts :
h, = £_ + St.. = t* u + 4.42 d. . (3-58)
A J *f j e L.
Trajectory - Reentrainment will depend upon gas velocity,
flow direction, terminal and initial velocities of drops and
hydrodynamic properties. Large drops which have a terminal
settling velocity greater than the upward gas velocity will
eventually fall down. However, it is necessary to provide
enough height following the mist eliminator, for these drops
to reverse direction and start falling.
3-57
-------
If the gas velocity is in the vertical direction, the
vertical height traveled by the large drops (with diameter
"ds £M) is given by Equation 3-59, from Lapple and Shepherd
(1940) :
U)
Udu
a6u
1'*
g
u du
a,ultlf - g
(3-59)
where
= -r^ - . The equation is based on drag coeffi-
cient Cd = 18.5 NR~°'6
u-u
The time for vertical travel is:
u,.
f du + f du_
I 1 U . I 1 U
J a,u • + e / a^u
* f\ & •?» A
(3-60)
Figures 3-22 and 3-23 give a plot of time and height
versus initial drop velocity with drop diameter as parameter
(gas velocity ur = 0).
The drop ballistics for small drops with drop diameter,
"d ", differs from that of the large ones because the
S y S
initial velocity of the small drops is not necessarily in
the vertical direction.
Lapple and Shepherd give integrated forms for this case,
based on Stokes Law:
PL dl.s
18 u-
Sh " ~IF
1 - exp
dl,s
(3-61)
(3-62)
18
U
- u
In
ut - u
Ut
PL
d2
1 - exp
'L us,s
(3-63)
(3-64)
3-58
-------
u
-------
where S » the distance traveled by drop and subscripts "h"
and "v" refer to the horizontal and the vertical
direction,
u = the initial velocity in horizontal (Equations
3-61 and 3-62) and in vertical (Equations 3-63
and 3-64) direction
Creeping of Fluids
The presence of drag forces due to gas flow leads to
creeping of liquid in the entrainment separator. Creeping
may be prevented by providing a proper drainage system. If
creeping is not prevented, reentrainment may occur.
In the entrainment separator gas usually flows vertically
upward or horizontally. If there is creeping of liquid, it
flows as a film or as drops. The liquid and gas flow in the
same direction. Thus, to determine the effect of creeping,
we need equations to predict liquid flow rate due to drag
forces exerted by gas in various directions and in various
geometries. Again, we are limited to the theoretical devel-
opment to predict two phase flow in simple geometries.
Consider liquid and gas flowing in a vertical tube. The
gas is flowing vertically upward and liquid is flowing as a
film and therefore forming an annulus. The liquid film is
subject to various forces: drag force due to gas flow in the
vertically upward direction, gravity force in the downward
direction and frictional force due to tube wall.
For gas velocity lower than the critical velocity the
liquid near the wall flows downward due to gravity. As the
gas velocity is increased the liquid at the interface reverses
its flow direction and moves with the gas; as a result the
liquid film begins to thicken. At a critical gas flow rate
the liquid does not flow down any more, and the liquid film
thickens rapidly.
3-60
-------
Shearer and Nedderman (1965) give the equation for shear
stress inside the liquid film:
dp _ _\ /D-2 6\ , / dp p
T =
at - PG g
• itn ' PL
C3-65)
where D = the diameter of the tube, cm
T = shear stress measured at a distance My" from the
wall, dyne/cm2
The velocity profile for the liquid can be obtained by
integrating Equation (3-65) . The liquid is assumed to be
Newtonian and the boundary condition is given by u. =0
LI
at y = 0 for a laminar flow:
u,
(3-66)
QT -
Integration of Equation (3-66) gives Equation (3-67):
gpT63 gprD62
\^IJ A/U . «LJ | U 1
' -& -5— + -tt I TT I '
L
irD
v
(3-67)
However, in the case of a liquid film flowing vertically
downward, terms containing "g" should be more important than
terms containing"-^",
az m „
The situation is significantly more complex in the
general case when the flow is not steady and part of the
film is in turbulent motion. In entrainment separators the
situation is further complicated due to entrance effect.
When the flow channel is not vertical, Equation (3-67)
should be modified by replacing "g" by "gCosB"; where "6" is
the angle between the channel direction and the vertical.
However, if the effect of gravity is larger than the effect
of the pressure gradient, the thickness of the film across
the channel will be uneven and separated flow rather than
annular flow may form. Equations are not available for the
case of cross flow.
3-61
-------
Shattering of Drops
Reentrainment may take place due to shattering of drops
in two ways.
1. Due to splashing of drops on the solid surface,
2, Due to high relative velocity between gas and
liquid drops.
Garner et al, (1954) observed splashing in a 10 cm
glass evaporator. The size distribution of the shattered
drops showed a large variation. The drops were larger
than 200 ym. The splashing was observed in a gas-liquid
system where localized areas of high velocity were present.
Shattering of the drops due to high relative velocity
between gas and liquid drops does not increase entrainment
in the gas phase. However, small drops are more liable to
be carried away in the gas phase than large drops and
therefore shattering of drops should be avoided.
MANUFACTURERS' SURVEY
The information received from the manufacturers of
entrainment separators was reviewed. Cyclone, mesh, fiber
bed, packed bed and baffles are commonly used as entrain-
ment separators. Sometimes two or more devices are combined
in the entrainment separation process. The description of
the available entrainment separators is given in the
proceeding paragraphs. The first paragraph gives the code
used in the description.
Manufacturer
(Type) Type of separator
(L) liquid load, cm3/cm3, and drop diameter, ym
(G) Gas volume, m3/sec
(UG) Gas velocity, cm/sec (maximum operating range)
(E) Overall separation efficiency, %
3-62
-------
(AP) Pressure drop, cm W.C.
(Lim.) Limitations
(App.) Application
(Mat.) Materials of construction
(Com.) Comments
ACS Industries
(Type) Mesh
(L) up to 2.71 g/cm2 sec, d > 5 ym
(G) UG = k [(PL - PG)/PG] k = 0.35 for oil and gas,
= 0.25 for steam at 1 atm, =0.2 for pressure
500 cm Hg, = 0.15 - 0.2 for vacuum
(E) Figure 3-24
(AP) Figures.3-25 and 3-26, Table 3-6
(App.) Mists of petroleum, H^SO., caustic, HC1,
water soluble dusts
(Mat.) . Steel, SS, MO, Ni, Haste alloy, PP, Teflon
Aerodyne Development Corporation
(Type) Cyclones, integral part of wet scrubber
(G) 14.1 (30,000 CFM)
(Com.) Uses secondary air to impart centrifugal action
AgetManufacturing Company
(Type) Fiber filter
(G) 0,141 - 1.22 (300 - 2,600 CFM)
(App.) Mists from screw machine, thread grinders,
centerless machines
(Mat.) Glass wool
Air Purification Methods
(Type) Cyclone with vanes
(L) 2.66x10-* -l.lxlO'3
(G) 0.19 - 15.5 (400-33,800 CFM)
(UG) 1,130 - 1,480 cm/sec
(E) 95% for 10 ym and 1001 for > 13 ym,
See Figure 3-27
3-63
-------
100
w
t—I
(_)
I—I
V*
O
U
20 SffigiS
0
0
2.5
GAS VELOCITY, m/sec
Figure 3-24 - Calculated collection efficiency
for water droplets in air for ACS mesh
3-64
-------
100
u
E
u
a,
o
w
CO
w
5 10
AIR VELOCITY, m/sec
50 100
Figure 3-25 - Pressure drop versus air velocity
for 10 cm thick ACS style 4CA mesh.
L/A = superficial liquid velocity,
cm/min.
3-65
-------
30
u
*
s
g
!=>
CO
0.5 -,
0.3
STEAM VELOCITY, m/sec
Figure 3-26 -
Pressure drop versus steam velocity at
various pressures through 10 cm thick
ACS mesh.
3-66
-------
Table 3-6. MESH PARAMETERS
Mesh Density Wire Diameter Percent Wire Surface
Style g/cm3 cm Voids cm2/cm3
4BA 0.192 0.028 97.6 3.77
4CA 0.144 0.028 98.2 2.79
7CA 0.080 - 0.028 99.0 1.48
3BF 0.115 0.015 98.6 3.93
3-67
-------
100
I i r
i i i i
10 15 20 25 30 35
PARTICLE DIAMETER, ym
40
Figure 3-27 - Collection efficiency versus
particle diameter for Air
Purification Methods cyclone
separator with inlet velocity
of 16-21 m/sec
3-68
-------
Air Purification Methods (continued)
(App.) in venturi scrubbers and mist removal
Arco Scrubbers (Envirotech)
(Type) Baffles, packed bed and cyclone
(G) 65 (138,000 CFM)
(Com.) cross flow in packed bed
B. B. Barefoot 5 Associates
(Type) Cyclone, integral part of scrubber
(G) 2.4 - 47.2 (5,000 to 100,000 SCFM)
(App.) Foundry, furnaces, sinter plants
Balston, Inc.
(Type) Fiber filters
(G) 0.38 (800 CFM)
(E) 99.99
(AP) 351
(App.) Sterilization of air, air to instruments,
laboratory air clean up, breather filters
in underground mines
(Mat.) Glass fiber 90% void
Beco Engineering Co.
(Type) Packed beds of 1.9 cm x 1.9 cm hexagonal
cylindrical pellets from three dimensional
bonded fibers
(L) No limit
(G) 35.2 (75,000 CFM)
(AP) See Figure 3-28
(App.) Distillation, fume scrubbers, packed towers
(Mat.) Nylon 6.6 for pH > 4, epoxy bonded polyester
for acidic medium
Beetle Plastics
(Type) Packed bed
(G) 0.16 - 33.6 (350 - 50,000 SCFM)
(UG) 220
(App.) Hcl, HF and NH3 fumes
3-69
-------
50
e
o
(X.
o
w
OS
D
CO
C/D
10.
5.0
1.0
0.5
I r
/Conventional
'pad
E-pack
Type E/N-1
I III il i i i I
1.0
10
AIR VELOCITY, m/sec
Figure 3-28 - Comparison of pressure drop
characteristics of Beco Engineering
Type E/N-1 pad and conventional
Pad length is 30 cm and air at room
temperature and 1 atm pressure.
3-70
-------
Burgess Industries (Delta P Division)
(Type)
U)
(E)
CAP)
(Lim.)
(App . )
(Mat.)
(Com.)
(Type)
(L)
(Mat.)
(Type)
(L)
(E)
(Mat.)
Cal-Metex
(Type)
(UG)
(E)
(AP)
(Lim.)
(Mat.)
(Com.)
Centrifugal
>5 vim, liquid load less than 51 by weight
99+% See Figure 3-29
<1% of abs. line pressure
Maximum pressure 25 atm
Compressors, expanders, vacuum pumps
Carbon steel
Coalescer recommended for below 5 ym
Small cyclone tubes
Maximum 5% load
95% for 4 ym, 98% for 6 ym
Carbon steel, stainless steel and alloy steel
Zigzag baffles
liquid load < 5%, drops > 5 ym
99.991 for 8 ym
Carbon steel
Mesh
30-460, 10% higher for hi-thruput
99.9+%
= 0.034 pG u* (Figure 3-30) for 10 cm thick
standard mesh, - 0.02 PG u5, for 15 cm thick
hi-thruput mesh
Sauder Brown constant, a = 0.25 for standard,
= 0.23 for hi-thruput and 0.22 for extra dense
S.S., carbon steel, Ni, Cu, PP
density = 0.145 g/cm3 for standard, = 0.08 g/cm3
for hi-thruput and = 0.195 g/cm3 for extradense.
See Figure 3-31 for onset of reentrainment
velocities.
3-71
-------
I
55 -6
° 10
o
OH"*
n e
H -
O O
i—i
«H FH
U
10
-7
10
-8
Figure 3-29.- Efficiency curve for Burgess
Industries centrifuge. Inlet
drop diameter is greater than
5 urn and maximum liquid in
inlet is 5% of gas weight.
3-72
-------
10
u
OH
o
en
I
in
w
0.01
I I I
-Nu-Standard
1.0 2 3 4 5
GAS VELOCITY, m/sec
Figure 3-30 - Pressure drop characteristic
of 10 cm-thick Nu-standard
and 15 cm thick Hi-thruput mist
eliminators.
3-73
-------
5.5
u
0)
S 5.0
>: 4.5
H
I 4-°
>
7 5
t/j ** • **
3.0
Hi-Thruput
— Nu-Standard
I
0 1 2 3
ENTRAINMENT LOADING, g/cm2-min
Figure 3-31 - Safe operating limits for
Nu-standard and Hirthruput
mist eliminators. Reentrainment
will not occur when velocities
do not exceed those shown.
3-74
-------
Cebeco Manufacturing Company
(Type)
CL)
(G)
(AP)
(App . )
(Mat.)
(Com.)
Ceilcote
(Type)
(L)
(G)
(UG)
(E)
(AP)
(Lim.)
(Mat.)
(Type)
(L)
(G)
(UG)
(E)
(AP)
Fiber filter
0.5 urn
118 (250,000 CFM)
99.9% for > 0.5 ym
17.5-25
H2SO^ mists, organic plasticizers, oil mists,
carbon black
Glass wool - finer than used in laboratories
5 cm thick, wash may be'used for solids
Packed column - crossflow and counterflow
>3 vim, >8 ym for countercurrent with upstream
washing
0.15 - 23.6 (330-50,000 CFM)
215 (maximum 255 - for lower liquid loads)
Crossflow: 95+% for > 10 ym and 85-95% for
<10 ym. Counterflow: 85-95% for > 10 ym and
50-85% for < 10 ym
0.008 cm/cm of packing at recommended velocity
in crossflow
If solids are present countercurrent flow is
recommended
Steel, steel lined with flakeline, rubber, etc
Wet cyclone
>5 ym
1,4-23.6 (3,000-50,000 ACFM)
1,650
95%
5-10
3-75
-------
Centri Spray Corporation
(Type) Zigzag baffles
(G)
CUG)
14.1 (30,000 CFM)
195
Scrubbers in foundry cupolos
Galvanized steel, PP
(App.)
(Mat.)
Chemico Industry
(Type) Cyclone
(Com.) Cyclone is integral part of scrubber
Donaldson Company
(Type) Multiple cyclone
(G) 26 (55,000 CFM)
(E) 99.97 for > 0.3 ym
(App.) Air purification in mines
(Com.) Irrigation with 2,66xlO'5 - 0.67xlO-s
L/G recommended
DuPont
(Type) Fibrous bed, 0.1-0.3 mm diameter fiber
(G) 14.1 (30,000 CFM)
(App,) Foundry cupolos, mist of H2S04, mists from
venturi scrubber
(Com.) Irrigation with water to avoid plugging,
flooding correlation shown in Figure 32
Edwards Engineering Corp.
(Type) Mesh
(L) Drops formed by condensation
(App.) Vapor condensers
Fisher Klosterman, Inc.
(Type) Cyclones
(L) Critical drop diameter 6-27 \im
(G) 19.6 (42,000 CFM)
(UG) 1,400
3-76
-------
O4
I
•vl
P-
id
i
O
0
0)
Q.
O
Q.
Figure 3-32 - Flooding correlation for fibrous bed structures by
Porter and Lucas (1968)
-------
Fisher Klosterman, Inc. (continued)
(E) 1001 above critical diameter drops
(AP) 15
(Type) Three stage filter consisting of 14-18 mesh,
fiberglass, submicron diameter fiber filter
(G) 15.1 (32,000 CFM)
(AP) 0.25 - 7.5, 0.26 at rated capacity
(Lim.) Filters need replacement
(App.) Oil mist collection
Industrial Plastic Fabricators
(Type) Packed bed 15 cm deep or mesh 10 cm deep
(G) 0.23 - 11.8 (500-25,000 CFM)
(E) 90-99% (efficiency of total scrubber)
(Mat.) Knitted mesh or PP
(Com.) Integral part of fume scrubber
Koch Engineering Company
(Type) Cyclone
(G) 1.18-47.2 (2,500 - 100,000 CFM)
(AP) 2.5-5
(Com.) Wall wash available if scaling is expected
(Type) Zigzag vane 3 or 6 passes
(UG) 355
(AP) 2.5 for 6 passes
(Mat.) PP
(Type) Spin vane
(G) 47.2 (100,000 CFM)
(AP) 10
(Com.) Less wear on shell, effective for vessels largef
than 400 cm in diameter
(Type) Fleximesh
(L) 2.7xlO"3 to 2.7xlO-2 g/sec cm2
(UG) 150-450
3-78
-------
Koch Engineering Company (continued)
(E) See Figure 3-33
(AP) 1.25 - 7.5 depending upon liquid load
(Lim.) Not recommended for high liquid loads (5x10 "''L/G)
(App.) In wet scrubbers, removal of acid mist, oil drop-
lets, in distillation columns
(Mat.) SS, Inconel, Monel, Ni, Ti, PP, Teflon, Hastelloy
Lau, Inc.
(Type) Centrifugal, mechanical aided and fiber filter
(AP) 12,5 maximum
(App.) Oil mist created by coolent or cutting lubricant
(Mat.) Polyurethene-filter
Monsanto-Enviro Chem Systems, Inc.
(Type) Fiber filter, high volume and high efficiency
filters
(L) No limitation on high efficiency, 3 pm on
high volume
(UG) <30 for high volume, <15 for high efficiency
(E) See Figure 3-34
(Lim.) Cannot handle excessive solids, upstream washing
recommended to avoid plugging
(Mat.) Fibers < 10 ym in diameter
(Com.) Above velocities represent onset of reentrainment
North American Rockwell
(Type) Mesh, panel filters, two stage on line mesh
separators
(G) up to 4.7 (10,000 CFM) maximum actual volume
= 0.47 (1,000'CFM) at 10 atm
(E) 99.96 on 0.3 urn drops - by DOP method
(AP) 60-250 when clean, 350-500 at 10 atm
(Lim.) 10 atm, 120°C
(App.) Removal of mist from compressed air
(Com.) All the data is for two stage mesh
3-79
-------
u
2
U4
M
U4
H
CJ
U4
246
GAS VELOCITY, m/sec
Figure 3-33 - Performance of Koch fleximesh
separator
3-80
-------
I
CO
u
2
C_J
I— I
U4
A
w
2
o
HH
H
u
W
100
99
95
90
85
80
75
70
(7)Brink H-E mist eliminate
0.3 0.5 1.0 2
PARTICLE SIZE, ym
10
20
Figure 3-34 - Collection efficiency of particulate collectors.
-------
Research Cottrell
(Type) Cyclone - integrated with flooded disc wet
scrubber
(L) 8-20
(AP) 2.5
(App,) cupolas, electric arc furnace, gas.cooling
(Mat.) 316 SS
Tailor § Company
(Type) Cyclone
(G) 141 (300,000 ACFM)
(E) exit gas < 2.33xlO"6 L/G
(Com.) Cyclone connected to sicrubber
Wright Austin Company
(Type) Cyclone
(G) 0.0025-47 (5-100,000 SCFM)
(UG) 4,050 no reentrainment up to 7,800
(E) 99* for > 10 ym
(AP) 50
(App.) Removal of drops from wet steam, compressed
air and gas
(Com.) Can be directly connected to steam lines
0.6 cm to 120 cm in diameter
York Separators
(Type) Mesh single stage and two stage
(L) 5 ym for single stage, 0.3 ym for two stage
(G) 450-540 (950,000-1,040,000 CFM) {
(UG) Sauder Brown constant for single stage, k, s 0'
for 7.5 cm disengaging space and k = 0.43 for
38 cm disengaging space
(E) 99.9 for single stage, 100% up to 1 ym for
2 stage - See Figure 3-35
3-82
-------
0\°
100
u
w
S 90
pt,
PH
W
z 80
o
H
• 1 «
York Type S
M_
—
U
W 7nL_
j 70h-
o 1 1 1 1
U 123
1 1
/Two Stage
Wire
Mesh
-
j 1
4 5
PARTICLE SIZE, ym
Figure 3-35- Efficiency comparison of
York scrubber with two stage
wire mesh mist eliminator.
3-83
-------
York Separators (continued)
CAP) <2.5 for single stage, 25 for two stage
(App.) Refinery vacuum towers, distillation equipment,
evaporators, absorbers, scrubbers
(Mat.) SS, Mo, Ni, Inconel, Ti, Ta, Cu, Hastelloy, PE,
PP, Teflon
Zurn Industry
(Type) Impingement separator
(G) 0.57 to 34 (1,200 to 72,000 CFM)
(type) Cyclone - integral part of venturi scrubber
(G) 56.3 (120,000 CFM)
3-84
-------
THEORY
In the previous chapter equations predicting primary
collection efficiency and pressure drop were presented
for cyclone, packed bed, tube bank, mesh and sieve plate
entrainment separators. Part of this chapter is devoted
to the development of mathematical models for primary
efficiency and pressure drop in a zigzag baffle separator.
In addition, models for reentrainment in both horizontal
and vertical baffles and in cyclones are discussed. At
this point in the study, these are the only separators
for which mathematical models of reentrainment have been
formulated.
ZIGZAG BAFFLES
Primary Efficiency
A zigzag baffle section is shown in Figure 4-1, The
gas flow pattern in a baffle section is too intricate
to be described as a series of gentle bends. The model
used to characterize the flow must include the effects
of turbulent mixing. One can assume either that mixing
occurs after flow through each bend or that mixing occurs
continuously throughout the flow path. The latter assumption
is made here.
The presumption of turbulence may be verified by
calculating the gas Reynolds number as follows:
Hydraulic Diameter - 4* - ' "> - C4-1)
2b'uipr 2burp
4-1
-------
n=4
t
u.
Figure 4-1 - Continuous zigzag baffles
4-2
-------
where b1 » perpendicular distance between two consecutive
baffles in the same row, cm
u,!, = actual gas velocity between the baffles, cm/sec
b = spacing between two consecutive baffles in the
same row, cm
ufi = superficial gas velocity based on empty duct,
cm/sec = uiCosS
Under typical conditions used in the pilot plant,
b = 7.25 cm, pfi = 1.25 x 10'3 g/cm3, yfi = 1.8 x 10-* poise
and UG « 300 cm/sec. Therefore, NRe G = 3 x 10" and the
flow should be turbulent. The presence of corners and edges
in the baffle section should enhance the turbulence.
The gas is assumed to flow as in Figure 4-1. The flow
path is a series of alternate bends of angle 26, and the
total number of bends equals the number of rows of baffles
in the entrainment separator.
The terminal settling velocity of a drop, in cm/sec, is:
d 2a a
ut - -f-2- (4-3)
where d «= drop diameter, cm
p = drop density, g/cm3
UG - gas viscosity, centipoise
a = acceleration due to centrifugal force, cm/sec2
The acceleration due to centrifugal force was approxi-
mated on the basis of an average radius of curvature in the
baffles, as follows:
(4-4)
w cote
Equation (4-3) is based on Stokes law and applies only if
the drop Reynolds number NRe D < 0.1. If NRg D > 0.1, an
appropriate friction factor should be included in the
4-3
-------
calculation. Foust, et al. (1959) give a plot of drag co-
efficient as a function of Reynolds number in Figure 4-2,
which can be used to determine "ut". The effect of sur-
rounding drops on the motion of any individual drop is
neglected.
A number balance on the drops entering and leaving an
element of fluid rd6 long, where r ~ cot 0 (see Figure 4-1),
Lt
b1 thick and of unit width gives:
u^b'dc = cutrde (4-5)
where c » number concentration, #/cm3
For the boundary conditions c » c (initial concentration)
at 6 *» 0 and c e c at 9 = 2n0 (where n B number of rows),
Equation (4-4) can be integrated to yield:
S- = Pt = exp --J ±gji - exp -ji jSgjj C4-6J
Ob b
where Pt = penetration
Figure 4-3 shows penetration versus gas velocity for
n = 6 with "9" and "d " as parameters. For a single baffle
angle penetration is predicted to increase as the drop dia-
meter is lowered. Increasing the angle should result in
decreased penetration for a particular drop diameter. In
all cases increasing gas velocity is expected to lower the
penetration. The effect becomes more pronounced as "9" and
"d " are increased.
If drop mixing is neglected, penetration can be
expressed as: u
Pt - 1 - — nw9- (4-7)
Ft L UG b tane ^ /J
A comparison of the theoretical primary efficiency curves
is presented in Figure 4-4. The model based on mixing gives
4-4
-------
10,00
o
h
w
(D
•H
o
(1)
(U
O
o
03
FH
P
0.001
0,01
D vp
Reynolds number (N,, =-2—1
<->n y J
Re
Figure 4- 2 -
Drag coefficient versus Reynolds number after
Foust et al (1959), with sphericity
the parameter.
as
4-5
-------
Z
O
h-1
H
H
w
PL,
= 30° d, = 30 ym——i-
3.0 4.2 5.4
VELOCITY, m/sec
Figure 4-3 - Penetration versus gas velocity
for baffle section.
4-6
-------
u
Z
W
hH
U
2
O
I-H
s
1-J
O
§
Q
100
80
60
40
20
30 urn ortri
20 ym
45°, dd = 10 urn
1.8
2.4 3.0 3.6 4.2
GAS VELOCITY, m/sec
4.8:
Figure 4-4 - Comparison of primary efficiency curves
based on theoretical model for complete
mixing ( ) and for no mixing C ) models
4-7
-------
an asymptotic increase to 100% efficiency as gas velocity
is increased, while the model based on no mixing predicts
a linear increase. However, experimental values of primary
efficiency in entrainment separators do not show a linear
increase with increasing gas velocity. Hence the model
based on mixing is considered more valid.
This model was developed for a continuous baffle
system. If the baffles are discontinuous and staggered,
the efficiency will be higher than that predicted from the
present model.
Pressure Drop
Determination of the pressure drop is based on the
drag coefficient, "fD", for a single plate held at an
angle "6" to the flow as presented in Figure 4-5 (Page
and Johanson, 1927). Neglecting the effect of neighboring
plates, pressure drop may be expressed as:
n u'2
AP ^Z £D pG -|- (4-8)
where the summation is made over the number of rows of
baffles.
The actual gas velocity,"uA", in the baffle section
should be used in Equation (4-8). Note that the angle of
incidence for the second and subsequent rows of baffles
will be twice the angle of incidence for the baffles in
the first row.
Reentrainment
The following sections are devoted to a theoretical
development of equations describing the conditions necessary
for the occurrence of reentrainment in baffle-type entrain-
ment separators. The first section deals with reentrainment
4-8
-------
1.2F
0.8
M-l
U
i— i
IX,
PH
Plate Inclined to Flow
: : . r ::: I . : ri
0
0
20 40 60
ANGLE OF INCIDENCE
80
Figure 4-5 - Drag coefficients for flow past
inclined flat plates (data from
A. Page $ F. C. Johansen, (1927)
4-9
-------
in horizontal baffles, and the second with the same
phenomenon in vertical baffles.
The significant difference between vertical and
horizontal baffles is that in horizontal baffles:
1. Prior to the onset of reentrainment conditions,
liquid flows on both sides of the baffles.
2. The wetted perimeter for gravity flow is in
proportion to the length of the baffle.
3. Depending upon baffle angle to vertical, only
the "g cos 6" component of "g" is available for
flow due to the gravity.
4. All the liquid has to flow through the bottom
row of baffles.
5. The drag force due to gas velocity and the
force due to the gravity are acting in opposite
directions.
Whereas in the vertical baffles:
1. Liquid flows on only one side of the baffle.
2. The wetted perimeter for gravity flow is pro-
portional to the width of the baffle.
3. Gravity is acting in the optimum direction for
liquid collection.
4. Liquid can be collected on each baffle.
5. The drag force due to gas velocity and the
force due to gravity are acting at right angles.
The effects of corners, capillaries and film detach-
ment have not been accounted for. Thus, it is possible
for reentrainment to occur before the conditions the model
predicts. Nevertheless, the model can be used to determine
the upper limit of operating conditions.
4-10
-------
Horizontal Baffles - The goal of this section is to derive
a relationship which gives the theoretical maximum liquid
load, "QT/QQ", at a given air velocity, "UG", with no re-
entrainment. The gas flow rate, "QG", can be expressed as:
A cm3/sec (4-9)
where a - baffle length, cm
The liquid flow rate, "QL", cm3/sec, is made up of flow due
to drag force between the gas and the liquid film and flow
due to gravity. The derivation of MQL" which follows,
results in a system of six equations in nine variables. At
that point two additional equations will be formulated in
order to determine the desired relationship. For certain
selected values of "UG", the maximum "QL/QG" cai* then be
found. A diagram of the results is presented.
Derivation of Q. - The following assumptions are made in
considering liquid flow in horizontal baffles:
1. Prior to liquid reentrainment, circulation of
the liquid on the baffles takes place, i.e.,
both sides of the baffle are wet.
2. Interfacial stress is the same on both sides of
the baffle, i.e., film thickness is the same on
both sides of the baffle.
3. Film flow takes place on the baffle.
The derivation begins by considering the forces acting
on a liquid film of thickness, "6", as shown in Figure 4-6 54-7.
A force balance on the liquid film bounded by the planes
at a distance, "6" and "Y", frtfm the baffle surface gives:
(6-Y) dzprgcos9 + (P +
L az
•*• P(6-Y)£ (4-10)
4-11
-------
Figure 4-6 -
Forces on an element of
liquid film on baffle
S
cosG
u,
Figure 4-7 -
Effect of baffle edges
on reentrainment
4-12
-------
2
where T = shear stress, dyne/cm
Y = distance from baffle to any point in the
liquid film, cm
cm
g - acceleration of gravity,
sec2
P = pressure, dyne/cm2
T. = interfacial shear stress, dyne/cm2
Equation (4-10) can be rearranged to:
dp
T - T. - (6-Y)(pLgcose + 2|.) (4-11)
A force balance on the gas core gives:
2T-A - -(b'-2<5)£ || (4-12)
For laminar flow in the liquid layer:
T = -U ~ (4-13)
Substitute the value of "TJ" from Equation (4-12) and
the value of "T" from Equation (4-13) into Equation (4-11)
and rearrange to yield:
du = + (6.y)(p gcose + , dY (4
V L 2. dz L dz J
Integrate over the limits 0 to "u" and 0 to "Y"
(no slip flow) to obtain the velocity profile:
u .
4-13
-------
"QL" can now be determined by substitution into its
defining equation:
6
QT = 2y*u£dY (4-16)
which results in:
(pL,co,e*«)] C4-17)
Neglect the term containing "63-7=-" as compared to
dz
"62__" (to check: effect of neglecting = ,(,.. If
dz D / i
6 - 0.1 cm and b' = 5 cm, effect = °'1 x 2 < 2%) to obtain:
D X «J
Equation (4-18) expresses the liquid flow rate as the
sum of its pressure gradient and gravity components. How-
ever, it is valid only for laminar flow. Sherwood and
Pigford (1952) and Grimley (1948) give the following
conditions for determining the flow regime:
Laminar flow without rippling, NRg ^ < 4 to 25
Laminar flow with rippling, 4 to 25 < Np
-------
C4-19)
Therefore, the expression for total liquid flow rate with
rippling is:
Additional Equations - Having derived an equation for "QT",
we now have two equations, (4-9) and (4-20), in five un-
knowns, "QL", "Q ", "UG", "6", and lldZ". To narrow the
CLZ
gap, we return to Equation (4-12):
(b'-26) dP _ b'dP
Ti = -— 2 ar- "2 ar
if 6 B o at the gas-liquid interface. By Equation (4-39):
Substitute (4-22) into (4-21) and solve for — :
dz
d7 = " ST £i UG2 " " b cos 6 C4-23)
where b = b' cos 6 and UG • UQ cos 0.
Equation (4-23) introduces a sixth unknown, "£.",' the
interfacial friction factor. Equation (3-38) states:
fi = £G + 1'5 hr-"fc"Tv/*7l C4-24)
4-15
-------
The two new quantities which appear in Equation (4-24) ,
"NR p" and "f /-.", can be related by commonly available
graphs correlating them. "NRe p" can be expressed as:
d u' pr 2b ur pr
NRe G = q u -- iT— - (4'25)
KC , ij y ^ y G
At this point we have six equations in eight unknowns,
and we need an expression for 6. The pressure on the
liquid film due to gas velocity, assuming a drag coefficient
of 1 , is :
P = !££ (4.26)
where P, = pressure due to gas velocity, dyne/cm2
ur = reentrainment velocity, cm/sec
Since gas is flowing on both sides of the baffles, a
force balance on the liquid film yields:
or
U2 5 <* 2q cos e (4-28)
Gc pp
Now we have seven equations (4-19, 4-20, 4-23, 4-24,
4-25, 4-28) and the correlation between "NRe G" and "£G" in
nine unknowns (QL, QG, UG,
-------
10
Of
~*x
J
Of
•V
Q
D
C/
9=30°=Ba££le Angl
10
6.0 7.2 8.4 9.6
GAS VELOCITY, m/sec
10.8
Figure 4-8 Maximum liquid load, QL/QG versus
gas velocity in horizontal baffle
section.
4-17
-------
liquid flow on the baffles, "Qi11* will be derived. Combined
with other equations, we will be able to construct a diagram
for theoretical reentrainment.
Liquid Flow Due to Pressure Gradient - The following assump-
tions are made in the derivation of "Q, " :
1. Reentrainment takes place only on the edges of
the baffles.
2. Film flow takes place on the baffle surface.
3. Liquid collected per unit area of the baffle
surface is the same everywhere on a given baffle.
4. Laminar flow with rippling is present everywhere
on the baffle.
Consider Figures 4-6 and 4-7, which show a portion of
the baffle with liquid film thickness, "6". A force balance
on the film bounded by planes at a distance "<$" and "Y" from
the baffle surface gives:
+ P(5-Y)£ (4-29)
which can be rearranged to :
T = T. - (6-Y) jjZ (4-30)
Assume laminar flow:
T " -Pay (4-31).
Substitute and rearrange:
du - - [>Ti - ^-Y)ar]dY (4-32)
A force balance on the gas core gives:
2TiA£ « - (b'-26)AAdT| (4-33)
Substitute the value of TJ from Equation (4-21) into
Equation (4-32) :
du . + (,-« dV (4-34)
4-18
-------
Integrate over limits "u" from 0 to "u" and "Y" from 0 to
"Y" (no slip flow) :
1 b dP
u . y . (4Y . j C4.3S,
The volumetric flow rate of liquid is:
QL " i"6 u AA dY (4-36)
where the integrand must be a small section of the baffle
length, "AA", rather than the entire length "A". Inte-
gration yields:
(4-S7)
If the term containing "63-— " is neglected compared to "6
dz dz
0 _ AA b'62 dP
QL ' vT T— 37 04-38}
dP
Substitute the value of "-gg" from Equation (4-23) to obtain
the liquid flow rate in the horizontal direction:
C4-39)
Liquid Flow Due to Gravity - As in horizontal flow a section
of the baffle must be considered, rather than the entire
baffle. Ripple flow on a vertical surface of width "Aw" is
of the same form as Equation (4-19):
<4-40>
4-19
-------
The differences in Equation (4-40) from Equation (4-19) are
caused by flow on only one side of the baffle and 6=0 for
vertical flow, making cos 0-1.
The total flow rate is the sum of the individual flow
rates.
Additional Equations - Equations for "6" and "f." have been
detailed in the derivation for horizontal baffles. The
expressions to be used for this case are identical.
In order to solve the system of equations for vertical
baffles, a stepwise calculation on a digital computer must
be carried out. In this instance, intervals of Aw = w, w/2,
w/3 and w/4 were used with AA = A, A/2, A/3, A/4, respec-
tively. In the present calculations reentrainment was
determined using Aw = w/4 and AA = A/4.
Figures 4-9 and 4-10 show the results of these calcula-
tions, with gas velocity as the parameter. Both give re-
entrainment as a function of liquid loading. Figure 4-9
utilizes a baffle angle of 30°, while for Figure 4-10, 9 - 45°
A comparison of the two figures shows that reentrainment
is higher and starts at lower superficial gas velocities for
6 = 45°. This is due to the fact that at the same super-
ficial velocity, the actual velocity in the baffle section
with 9 = 45° is higher than the actual velocity in the
baffle section with 9 = 30°.
CYCLONE
Primary Efficiency
The following derivation gives primary efficiency of a
cyclone as a function of the average residence time of a
drop in the cyclone. The derivation is based on the fol-
lowing assumptions:
1. Drop motion is not influenced by neighboring drops.
2. There is no coalescence of liquid drops in the gas
phase.
4-20
-------
3x10
CX
w
cx
E-
h-l
X
2:
H
H
2
W
W
oi
:m/sec
2x10
1x10
Zigzag bafEle
0 = 30°
j^ Jr My^l :y^T 1
?!>r:-w*«f-d>^ru:
Figure 4-9 -
IxlO'3 2xlO~3
QL/QG COLLECTED ON BAFFLE, m3/m3
Predicted effect of liquid loading in inlet on
reentrainment from vertical baffles. Gas
velocity is the parameter.
3x10-
4-21
-------
3x10-
Zigzag Battle
9 = 45°
sec
iHTiilH^l^ 4.2
lxlQ-3 2xlO'3
QL/QG COLLECTED ON BAFFLE, m3/m:
3xlO'3
Figure 4-10 - Predicted effect of liquid loading in inlet on
reentrainment from vertical baffles. Gas
velocity is the parameter.
4-22
-------
3. The radial velocity of the gas is zero.
4. The tangential velocity component of the drop is
the same as that of the gas stream, i.e., there
is no slip in the tangential direction between
the drops and the gas.
5. The tangential velocity component is related to
the radial position by a modified form of the
equation for a free vortex in an ideal fluid.
u. rn = constant (4-41)
where u = tangential gas velocity, cm/sec
r - radius, cm
n = vortex exponent, defined by Equation (3-5).
6. The drag force, turbulent mixing, and drop
bouncing or reentrainment are sufficiently
prevalent to insure that a uniform concentration
of uncollected drops is maintained in the gas
flowing through any horizontal cross-section of
a cyclone, i.e., that back-mixing is complete.
7. Reentrainment takes place only in the cylindrical
part of the cyclone.
8. The gas velocity near the vortex wall inside the
cyclone is equal to the gas velocity in the
cyclone inlet.
9. The gas enters the exit core (below the cyclone
exit) only in the conical portion of the cyclone.
A force balance on a drop in the cyclone yields:
dr^lSudr 9 ,n 1 . 'f A
4-23
-------
From Equation (4-41) :
d \n
Utg r • ui Vr (4"43)
where u, = cyclone inlet velocity, cm/sec
d = cyclone diameter, cm
Neglecting the second order differential from Equation
(4-42) and combining Equations (4-42) and (4-43):
If the drop just reaches the cyclone wall in time "At",
then r = d /2 and from integration of Equation (4-44):
d
At = ,..??. f,,.. "j 1 1 - /|I
2n+2
(4-45)
Consider a horizontal cross-section of a cyclone, as
shown in Figure 4-11. In time "dt", all drops a distance
"dr" or less from the cyclone wall will move to the wall
and be collected. Meanwhile the drops will travel a distance
"rde" tangentially and "dh" vertically. The number of drops
removed, "dn.", will be:
C4-46)
where c = number concentraion of drops, #/cm3
The total number concentration of drops in the sector
from which drops are removed is:
ni = nr dc c dh (4-47)
4-24
-------
Figure 4-11 - Cross section of a cyclone.
4-25
-------
The fraction of drops removed in time "dt" is therefore:
dn,
ni
d dr - dr'
c
4dr
(4-48)
neglecting the second order differential.
In order to relate the fraction of drops collected to
the average residence time, it is necessary to express
Equation (4-48) in terms of time. This may be done through
Equation (4-44) :
dr
ar
2d
2 dc
T"
2x1+2)
(4-49)
Combining Equations (4-48) and (4-49) :
i+l
/
At
ni
dn,
n.
Can+!
dt (4-50)
Integrating up to the average residence time "t" we obtain
the equation for primary collection efficiency:
- 2
(n+l)At
(4-51)
Reentrainment
Onset of Reentrainment - There is great disagreement among
results for the onset of reentrainment obtained by different
investigators. This is indicative of the problem of defining
4-26
-------
the onset of reentrainment . Zhivaikin's (1962) equations
will be used here for the onset of reentrainment:
r> ooo\t""0«75 4 £ XT ..0.085
P = 29'2 NRe,L lf NRe,L * "T (4-52)
p . 10OV
L e>L
VL L
P = 43.2 VT 1>25 ND . > 2JL1
L Re,L - VL (4-54)
where P = -2- (4-55)
It is assumed that reentrainment does not take place
if the onset of reentrainment conditions are not met. The
liquid Reynolds number used to determine onset of reentrain-
ment is based on the liquid flowing in the film. It is
assumed that the diameter of the inner core of the vortex
is the same as the diameter of the exit pipe, d . Thus, the
hydraulic diameter is (d - d ).
G &
Drop Diameter of Reentrainment - The drop diameter of the
reentrained liquid has a size distribution which varies
with gas flow rate, liquid flow rate, fluid properties and
perhaps pipe diameter. The average drop diameter decreases
with increase in gas flow rate. On the other hand, the
liquid flow rate has only a weak and ambiguous effect.
When the gas velocity exceeds 6,000 cm/sec, high liquid
flow rate has no effect on the drop size distribution.
Rate of Reentrainment - It is believed that reentrainment
takes place due to penetration of liquid waves into the
turbulent zone of the gas. The amplitude of the waves
increases exponentially with liquid flow rate. Therefore,
reentrainment is assumed to take place in proportion to
exp(Kn NT, T) where "K, " is a constant,
J. K6 , L -L .
4-27
-------
The rate of reentrainment depends upon gas flow rate,
liquid flow rate and fluid properties. According to Anderson
et al. (1964), the rate of reentrainment is approximately 41
for NR , > 2,750 and is seen to increase slightly with
NT?O r (3.5% for NDo r = 3 x 10", 41 for N~ r - 1.6 x 105) .
Re, G v Re,G f Re,G '
Below NR L = 2,750, the only data available are for
NR , = 1,150, at which point reentrainment is 0.5%.
Reentrainment Calculation - In order to determine reentrain-
ment in a cyclone, a stepwise calculation was carried out
with the aid of a digital computer. Each step represented
30 cm of gas travel. The primary collection efficiency,
reentrainment velocity, reentrainment and secondary col-
lection were calculated, and a material balance was carried
out at the end of each step. Some of the results appear in
the following four figures.
Figure 4-12 shows the predicted reentrainment from a
cyclone as a function of gas velocity with inlet drop dia-
meter as the parameter. Reentrainment is seen to be insensi-
tive to drop diameter at velocities greater than about 37.5
m/sec.
Figure 4-13 gives the theoretical reentrainment as a
function of liquid loading. A gas velocity of 30 m/sec and
an inlet drop diameter of 50 ym was used to plot this curve.
As expected, increased liquid loading should increase the
reentrainment under the stated conditions.
Figure 4-14 is a comparison of the primary and overall
efficiencies of a cyclone. The effect of the drop size
distribution is seen to be small. The overall efficiency
curves are obtained by subtracting reentrainment from the
primary efficiency curve.
4-28
-------
x
w
H
Z
W
3x10-"
2x10
IxlO'"
y'lBi
~ : ^-t-j h - f - • , ^
TI^Liquid load concentration in T
Tinlet = QL/QG - 1.33 l/
'
(10 gal/1,000 cu.ft.)
. L...
8^T id, (drop diameter of
*^T~ inlet entrainment)
150 ym
100 ym
50 ym
LiZiE
" * J
n
n
n-
n
i
— -— '
^
t
i
j
.
;
i
. - . ' i
I/.: ._•::._
1
t
. ! .
,
— :
.
rrr-
•
• Ti
-•—
.- . „
' ' '
I
r—
15 20
INLET GAS VELOCITY, m/sec
Figure 4-12 -
Predicted effect of inlet gas ^velocity
on reentrainment in cyclone with
a=9.2 cm, b=28 cm, dc=127 cm, hg=132 cm,
and d =67 cm.
e
4-29
-------
4x10"'
m
s
ry
W
cr
g 2x10-"
H
2:
S
T
Drop diameter in the inlet
entrainment = 50 ym
I
I
IxlO'3
QT/Qr IN THE INLET, m3/m3
Ju Li
2xlO'3
Figure 4- 13 - Predicted effect of liquid
loading in inlet on
reentrainment in cyclone
with a=9.2 cm, b=28 cm, d =127 cm,
h =132 cm and d =67cm.
3 ^
4-30
-------
100
50
1
0
15
Primary efficiency (no reentrainment)
Overall efficiency (with reentrainment)
Reentrainment drop diameter = 200 ym
Reentrainment drop diameter - 125 ym, 250 jam'
(50% each)
Drop diameter in inlet gas = 50 ym
Liquid load in inlet gas, QL/QG - 1.33x10~3
30
45
60
INLET GAS VELOCITY, m/sec
Figure 4-14 - Predicted effect of inlet gas velocity
on efficiency of cyclone with a=9.2 cm,
b=28 cm, d =127 cm, h =132 cm and
de=67 cm. c s
4-31
-------
Figure 4-15 gives the predicted efficiency for an
efficient or long cyclone. A height of 381 cm for the
cylindrical part of the cyclone is used instead of 127 cm
as in Figure 4-14. Reentrainment is predicted to cause
a sharper drop in the overall efficiency for the longer
cyclone, despite the fact that overall efficiency is
higher.
4-32
-------
100
w
CJ
Primary efficiency (no reentrainment)
Overall efficiency
(with reentrainment)
Inlet liquid load = Q /Q
L G
30 45
INLET GAS VELOCITY, m/sec
Figure 4-15 -
Predicted collection efficiency with and
without reentrainment.for cyclone with
?;2'3 C^ Cm' dc=127 cm» ^,,=381 cm,
and d =67 cm. c s
w
-------
4-34
-------
AUXILIARY EXPERIMENT
As discussed in Chapter 3, the first mechanism for
reentrainment, i.e. transition from separated to separated-
entrained flow caused by high gas velocity, represents
the theoretical upper limit of entrainment separator
operation. However, in actual practice reentrainment
takes place at much lower velocities due to the fact that
gas and liquid do not flow in a simple geometry as in
annular flow. For instance, jets of the gas stream may
strike the water film at an angle, resulting in entrained
droplets.
In view of our need for additional information, a
small scale experiment was carried out to make visual
observations of the transition from separated flow to
separated-entrained flow, including the following:
A. Observe the effect of duct dimensions.
B. Study entrainment velocities as a function of the
liquid flow rate.
C. Study drop size distribution created by the above
transition.
D. Study the effect of impingement angle between
air and water phase on entrainment velocities.
For design purposes it is necessary to know the effect
of duct dimensions and liquid flow rate on the transition
from separated to entrained flow. The drop size created
by reentrainment will govern the efficiency of separation
ff
of any later stages. For example, a series of two entrain-
ment separators can be used so that the first stage acts as
a high-velocity coalescer, while the second is a low-velocity,
efficient separator. Thus, the drop size of the reentrainment
in the first separator will play an important role in the
overall efficiency.
Cousins and Hewitt (1968) and Wicks and Dukler (1966)
5-1
-------
studied the drop size distribution obtained in the separated-
entrained flow. Chien and Ibele (1962), Steen and Wallis
(1964) and Zhivaikin (1962) studied the onset of entrainment
in separated flow. Most of this work was done on small
diameter tubes and there are some differences in the results,
Thus, it is necessary to check their results and obtain
further information for design of entrainment separators.
No work has been reported on the effect of impingement
angle.
Figure 5-1 shows the apparatus used to study two-phase
flow. It consisted of an open channel 6.5 cm wide and
inclined at 30° from the horizontal connected to an over-
flow tank at the top and a collection funnel at the bottom.
The channel was 77 cm long and made of aluminum. Air was
blown through a nozzle 5.4 cm x 1.9 cm directly on the water
film in the channel at various angles from 0° (parallel to
water flow) to 40°. A pitot tube connected to an inclined
manometer was used to measure air velocity.
The liquid was introduced into the channel from the
overflow tank. Air was introduced with the air inlet
nozzle positioned slightly above the liquid film so as
not to disturb it. Visual observation and a technique
utilizing chemically treated filter paper were used to
determine the transition from separated to separated-
entrained flow.
SAMPLING FOR DROPLET SIZE DISTRIBUTION
Due to its simplicity, the chemically treated filter
paper technique was used to obtain drop size distribution
of the entrainment. The method is described by Chilton
(1952) and utilizes Whatman filter paper #1 treated with
alkali fast green for sampling. The blot diameter is
5-2
-------
en
t
Pitot
Tube
Blowe
Air Nozzle
Open Channel
Figure 5-1 - Apparatus to study onset of entrainment velocities
-------
usually larger than the drop diameter due to expansion, and
the correlation between the diameters is given by Chilton.
It has been determined that there is no effect of using
different chemicals,except that chemicals affecting surface
tension will have a different correlation.
Some errors in the sampling of drop diameters by
filter paper are discussed below:
(1) If the concentration of the blots on the filter
paper is high, a few of the blots will overlap. A correct!0"
is necessary if the area occupied by the blots exceeds 10$
of the area of the filter paper.
During the sampling, drops spread in the filter paper.
It is assumed that the blots in the filter paper all have
thickness "t^" which is independent of the blot size in the
range of diameters measured on filter paper. Then, by
material balance:
7r^2i--7r^3 re -n
T db tl " 6" dp C5'1)
where, t, = blot thickness, cm
d, = blot diameter, cm, if blots do not overlap
:. d a db2/3 (5-2)
The 2/3 slope in the correlation between blot diameter
and drop diameter by Chilton (1952) confirms Equation (5-2)-
Thus, even if the drops overlap, the area occupied by the
blots on the filter paper will be the same.
Sauter (1952) showed by statistical analysis that if
drops in air are projected onto a plane, the measured
value of the fractional area occupied by the shadows, "A1",
is related to the actual value of fractional area, "A", by:
A = -ln(l-A') (5-3)
5-4
-------
where A = the fractional area of the shadows if the drops
do not overlap. Thus, in the present case:
A a Nd£2 (5-4)
and A' a N'd£2 (5-5)
where N' = number of blots in a given area
N = number of blots in a given area if blots do
not overlap.
d/ = measured blot diameter, cm
b
- N = N'jL (5-6)
N - N' (rAnr) (5-7)
Since, Ndb2 = N'd^2: (5-8)
db = diJ±-T- C5-9)
This is the average correction and should be applied
to the mean diameter of the blots.
The fractional area can be obtained by simply measuring
the blot diameters in a unit area and using the following
relationship:
A = I I N' d,;2 (5-10)
(2) While measuring the blot diameters, care should be
taken not to repeat or leave out some blots. This error
possibility may be reduced by using the same technique used
to count blood cells.
(3) The capture efficiency of the filter paper is
not 100%. A correction is necessary to obtain the actual
size distribution.
5-5
-------
To determine the error in the sampling, Figure 5-2 was
constructed. Drop collection efficiency of a filter paper
is plotted against drop diameter, with gas velocity as
a parameter. Drop collection efficiency was obtained from
the correlation between impaction efficiency and inertial
parameter by Golovin and Putnam (1962). The adhesion of
the particle to the collector is assumed to be 1001 in
this case. As seen in Figure 5-2, the error in measuring
drop size distribution at the gas velocities and drop
diameters used for these experiments is quite small.
DISCUSSION
The liquid film was about 1-3 mm thick and was
quite smooth, with no pulsation except when the entrainmen*
velocity was approached. The liquid film thickness was
constant in the bottom of the channel in the absence of the
air flow. With air flowing, the liquid film thickness in
the corners of the channel was higher than in the center of
the channel. This increase in the flow rate of the liquid
in the corners increases with increase in the total liquid
flow rate.
The liquid film thickness at the point of air impinge^6
was much smaller than the thickness upstream from that p
(for impingement angle >0°). At a certain distance
downstream from the point of air impingement, the liquid
film regained its former thickness. Entrainment was
observed to take place near the downstream spot where the
liquid film rose.
The air velocity measured is the actual air velocity
and should be distinguished from the superficial air
velocity used by other investigators. If the liquid film
1-3 mm thick and the tube diameter 2.5 cm, the difference
between actual and superficial velocities may be 10 to 40$
5-6
-------
100
0\°
v—'
u
w
t-H
U
I—I
w
o
U
W
O
U
O
rt
Q
100
DROP WAMETER, urn
Figure 5-2 - Predicted effect of drop diameter on
capture efficiency by 5 cm diameter
filter paper held perpendicular to
the air velocity.
5-7
-------
RESULTS
Figure 5-3 shows the drop size distribution in the
separated-entrained flow. The drop size distribution was
measured by increasing the air velocity to the point where
measurable entrainment was first observed on the coated
filter paper. There is no significant effect of different
flow rates on the drop size distribution. The average drop
diameter is about 250 ym. The distribution of the drop
sizes is such that the majority of the drops are small, but
these contribute relatively little to the total volumetric
drop flow.
Figure 5-4 represents, on a log-probability diagram,
the same size distribution as shown in Figure 5-3. It is
log normal with mass median drop diameter of 250 urn and
geometric standard deviation of 1.5.
A comparison of drop size distributions obtained by
various investigators with the present results is shown in
Figure 5-5. The data of Cousins and Hewitt (1968) were
obtained during experiments on the mass transfer of liquid
droplets subsequent to the removal of the liquid film.
Photographs of the droplets were obtained for a zone imme-
diately adjacent to the channel wall, and the images on the
photographs were measured to produce drop size data.
The Wicks and Dukler (1966) data were obtained by a
technique using variable spaced needles. The liquid was
injected using a porous wall section. These data should be
treated with some reservations, since the reliability of the
method of measurement is still uncertain. Figure 5-5 shows
that all the distributions are similar in form with slight
variation from present results near the end of the curve.
In the present runs the effect of gas velocity on the
drop size distribution was found to be negligible. However,
the data of other investigators, as shown in Figure 5-5,
5-8
-------
0
100
VOLUME PERCENT OVERSIZE
Figure 5-3 - Distribution of droplet size in air-water
entrained-separated flow.
5-9
-------
en
i
1,000
Pi
a,
o
10
20
30 40 50 60 70 80
VOLUME PERCENT OVERSIZE
90
95
98
Figure 5-4 -
entrained-separated
-------
2,000
1,000 -
Es
w
H
W
§
i—*
Q
a,
o
Present Investigation
Data of Cousins and
Hewitt (1968)
Data of Wicks and
Dukler (1966)
Liquid Pipe
Flow or
cm3/sec Duct
Air
Flow
Curve cm/sec
6.5 cm wide
open channel
107-191
Range ot |
4,900-8,500
flow rates
0.95 cm tube
15.2 Gmx
1.9 cm
vertical
duct
8,325
2,345
3,750
30 -
20
VOLUME PERCENT OVERSIZE
100
Figure 5-5 - Comparison of drop size by different investigators
5-11
-------
indicate that increasing gas velocity tends to reduce the
mass median diameter. In the velocity range depicted in
Figure 5-5, the mass median diameter decreases by roughly
40% when the gas velocity is doubled.
The effect of impingement angle between the gas and
the liquid stream is shown in Figure 5-6. It shows the
drop diameter versus volume percent oversize in separated-
entrained flow. Impingement angle is used as a parameter
and was varied for a given water flow rate until entrainment
was detected. These data were obtained with air velocity
of 38 m/sec through the nozzle outlet, a velocity smaller
than the critical velocity needed for entrainment in
parallel flow. Thus, the effect of impingement angle O0°)
is a reduction in entrainment velocity. From Figure 5-6 a
slight effect of impingement angle on the drop size dis-
tribution is seen.
The effect of liquid flow rate on the impingement angle
needed to produce entrainment is shown in Figure 5-7. Air
velocity was 38.8 m/sec during these runs. The effect of
increasing the impingement angle is to reduce the liquid
velocity required for entrainment.
Figure 5-8 compares the results for entrainment velo-
cities obtained by various investigators. Here UG<, = gas
velocity causing entrainment and the liquid Reynolds number,
"ND T", is the superficial liquid Reynolds number based on
K6 , L
the empty cross-section of the duct:
%a,L '
where VL = kinematic viscosity of the liquid, cm2/sec
d = duct width, cm
5-12
-------
E-H
W
8
Q
300
dNP_ A
200 —
100
70
60
50
40
30
20
Air velocity 38.8 m/sec
Impingement Water
Angle Flow
degrees cm3/sec
o
A
D
40
30
17
107
145
191
i
1
0
20 40 60
VOLUME PERCENT OVERSIZE
80
100
Figure 5-6 - Effect of impingement angle on drop size
5-13
-------
PJ
CJ
20
Air velocity 38.8 m/sec
10
1 1 1 1 1 1 1 1
1
100
120 140 160
LIQUID FLOW RATE, cm3/sec
180
200
Figure 5-7
Effect of liquid flow rate on impingement
angle for onset of entrapment.
5-14
-------
100
o
-------
The results differ significantly due to the problem of
defining the onset of entrainment. Zhivaikin (1962), who
defined the onset of entrainment as occurring when it is
first detectable, expressed his results in terms of the
following empirical equations for the critical gas velo-
city "UQ " for the onset of entrainment:
P = 29.2 N-°;[5 if NRejL<_°i (5.12)
4_l » A-.~ , .u V T
VL L
P = 43.2 vi"'. if NRe H^- (S.14)
where "P" is defined by:
p = l!§c^L (5.15)
Steen and Wallis (1964) defined the onset of entrain-
ment as that air velocity which represents the extrapolation
of the straight line portion of a graph of entrainment per-
centage versus air velocity. Since the increase in en-
trainment with air velocity is similar to the exponential
function, their results lie considerably above those of
Zhivaikin. Steen and Wallis also obtained an expression
for "P", when viscous forces in the liquid film can be
ignored, as:
P = 2.46 x 10-* (yL/uG)(pL/pG)^ (5-16)
Chien and Ibele (1962) defined the transition on the
basis of pressure drop versus gas flow rate curves. A
change in the slope of the curve was taken to indicate
the onset of gross entrainment. Their results apply to
5-16
-------
conditions where a large fraction of the total liquid flow
is entrained, so their prediction of "UQC" is even higher
than that of Steen and Wallis. Chien and Ibele correlated
their data by the following equation:
where NR G = gas phase Reynolds number with UG = uf
Equation (5-17) can be rearranged as:
1.2 x 10'
o
-0 . 3
(5-18)
The present results are comparable with those of Chien
and Ibele. In this study we defined the onset of entrain-
ment as when the entrained drops made an impression on the
treated filter paper. The present results may be slightly
higher due to the fact that actual air velocity in the
inlet nozzle rather than the superficial velocity in the
channel was measured.
CONCLUSIONS
A. Because our results were comparable with others
in terms of entrained drop diameter despite
differences in duct dimensions, it is concluded
that duct size does not affect the reentrained
drop size distribution.
B. Entrainment velocity depends upon liquid Reynolds
number as shown in Figure 5-8. The Chien and
Ibele curve is recommended for determining the
onset of reentrainment.
C. Increased impingement angles result in decreased
entrainment velocities as shown in Figure 5-7.
Thus, sharp angles in the entrainment separators
should be avoided.
5-17
-------
5-18
-------
EXPERIMENTAL PILOT PLANT
An experimental pilot plant for the study of entrain-
ment separators was designed and built. The purpose of the
pilot plant was to do the following:
1. To obtain reliable data over a wide range of
operating variables to provide a basis for
improved or new methods
2. To check presently available design equations
for entrainment separators
A. Efficiency of separation
B. Pressure drop
3. To determine the effect of higher gas velocity
on reentrainment, bouncing of drops and
impaction mechanism
4. To study liquid drainage and flooding
5. To study problems associated with entrainment
separators
6. To study vertical and horizontal sections
DESCRIPTION OF THE PILOT PLANT
The maximum capacity of the wet scrubber entrainment
separator is 85 m3/min (3,000 CFM). This capacity was
selected based on the following consideration. The entrain-
ment separator cross-section was selected to be 30.5 cm x 61 cm
This section is sufficiently large to have minimal wall effects
for separators and provides a fairly long (61 cm) collection
element when cross-flow effects are important. Normally, the
maximum air velocity in industrial separators is around
3.0 m/sec. If velocities 2,5 times higher are studied, the
maximum air velocity will be 7.5 m/sec. This will give the
maximum capacity of 85 m3/min.
6-1
-------
A block diagram of the pilot plant is shown in Figure 6-1.
The equipment consists of a filter, blower, heater, spray
sections, observation sections, entrainment separator test
sections, liquid tanks, pumps, etc. The top view of the pilot
plant is shown in Figure 6-2. Figure 6-3 shows the connections
between various tanks at the bottom of the platform.
A description of the equipment in the pilot plant is
given below.
Air Prefilter
Automotive air filters were used to clean the incoming
air. The air supply was taken from outside the building and
was connected to a box containing prefilters with a 48 cm
diameter galvanized duct. The prefilter consisted of 5 filters
in parallel, each having a capacity of 18.5 m3/min. The
pressure drop at this capacity is 0.25 cm W.C.
Blower
A Western Blower size 122 Bl and Class III was used. It
has a capacity rating of 88 m3/min at 30.5 cm W.C.(static
pressure), a 7.5 KW (10 HP) motor, rotatable housing and an
opposed blade discharge damper. The blower was supported on
a hard rubber base to prevent vibrations, and it had a flex-
ible duct connected at the outlet. The blower was insulated
with accoustical fiberglass and concrete blocks to reduce
the noise level.
Air Heater
A gas-fired air heater rated at 30 kcal/hr (120,000
Btu/hr) was used to heat the air.
Spray Section
The spray section served to humidify the incoming air
stream and to generate entrainment for the test section. It
6-2
-------
ON
OJ
Air
dry
wet
i j
dry
[•
wet - AP - wet
B! Obi
8. Liquid Catch Tanks
8
•Vent
Figure 6-1 - Block diagram of
experimental apparatus.
-------
Figure 6-2 - Top view of the entrainment separator pilot plant. All
dimensions in cm.
-------
i
Ol
Valves
Unions
Vent
— Flow
FP1
FP2
U
P2
Tl
ATI
AT 2
T2
-T6
- 110 fc/min
Pump
- 11 fc/min
Pump
- Vacuum Pump
- Recycle Pump
- Feed Tanks
- Catch Tanks
Figure 6-3 - Flow diagram showing connection between
various tanks and pumps.
-------
was equipped with various nozzles from Spraying Systems Co.
The nozzle specifications are given in Table 6-1. In any
section the nozzles were equispaced as shown in Figure 6-4
to generate uniform flow.
The 1/4 M6 nozzle, a single fluid hollow cone nozzle
with a 0.105 cm orifice, was operated at 3.4 atm line pressui"6
and produced drops of about 250 ym in diameter, in agreement
with the manufacturer's claim. If operated at 13.6 atm, it
is expected to produce drop diameters ranging from 40 to
100 ym.
The 1/4 M6SS nozzles were connected with 1.25 cm iron
pipe. The remaining nozzles were connected with 2.5 cm
p.v.c. pipe. Two spray sections were used in the pilot plant-
The 1/4 M6 nozzles were installed in the first spray section,
which had inside dimensions of 30.5 cm x 61 cm cross-section
and 46 cm length.
The second spray section, 76 cm in length, was equipped
with 1/4 M26 nozzles, and arrangement was made to accommodate
any of the remaining nozzles. There was considerable reducti0
in the cross-sectional area due to the presence of 2.5 cm
p.v.c. pipes. The reduction was compensated by enlarging the
cross-section to 46 cm x 61 cm. Whenever 1/4 M6 nozzles were
used to generate spray, the second spray section was replaced
by a blank section having a cross-section 30.5 cm x 61 cm.
The spray section was equipped with a drain at the bottom
to catch the liquid collected on the walls of the spray secti°p
Observation Sections
Two observation sections were used, one at each end of
the entrainment separator test section. They had dimensions
of 30.5 cm x 61 cm cross-section and 50 cm length. Two plexi'
glass windows, 30 cm x 30 cm, were installed on opposite sides
on each observation section. A door was provided for sampling
of entrainment and drop diameters.
6-6
-------
Table 6-1. NOZZLES USED IN SPRAY SECTION
Type of Nozzle
Hollow cone
Fogjet Nozzle
Full Cone
Hollow Cone
Model #
(Spray Systems)
1/4 M6SS
1-11 1/2 F18
1 1/2-11 1/2
F35
1/8 GG3
1/4 M26
Pressure
atm
13.6
2.7
2.7
2.7
2.7
# of
Nozzles
12
1
1
12
12
Flow rate
cm /sec
nozzle
14.2
1140
2200
63.0
27.2
-------
— c
fl
1
- 7.6 —
rs f
J A,
c
\ ' r
J
f
\.
V
5.3-*
V.
•\
J
f
\
j
•
5.1
fl - - im
f
k
k |
J
")
f]
!
—5.1
*\
— 7.6 —
to
*
CO
1
1
C1
0
a
3
r>
CO
c
a
c
a
c
•
1
r>
• cH
D VO
•^
3
f)
CO
4-
•
CO
1
Figure 6-4 - Nozzle positions in the 30.5cmx61cm
duct. All dimensions in cm.
6-8
-------
Drainage of Liquid in the Test Section
Liquid drainage is important in the design of entrain-
ment separators. The original drainage system in the test
section consisted of 20 equispaced openings, 0.6 cm in
diameter, per 30 cm length of the test section and a false
bottom, giving a space of 1.8 cm height, which was filled
with Raschig Rings. It was observed that the drainage in
the test section was not 100% effective. The bottom sur-
face was coated with hydrophobic resin, and the collected
liquid had a tendency to flow around the holes and thus
creep along the length of the test section.
To solve this problem the holes were enlarged to
0.8 cm diameter, and the number of holes was increased
from 20 to 40 per 30 cm length of the test section. To
avoid creeping of fluids at the bottom of the test section,
notches 0.6 cm wide and 28 cm long (perpendicular to the
air flow direction) were made at 30 cm intervals in the
test section bottom. The number of Raschig Rings was
reduced to about 1/3 under the false bottom. This drainage
system was quite efficient for low liquid loads.
The theoretical maximum air flow (without Raschig Rings
in the false bottom) through the false bottom is 21 of the
total flow. The presence of liquid and Raschig Rings in the
false bottom should lower this percentage significantly.
No reentrainment from the false bottom was observed. At
higher liquid flow rates (57 fc/min and higher), the liquid
built up to 3-4 mm in height for drainage. The total
drainage capacity of the test section was more than 150 Jl/min,
However, most of the liquid was collected in the first two
drainage sections of the test section.
The notch dimensions were increased to 2.5 cm width
since it was observed that a notch is more effective than
a hole in drainage. This change resulted in efficient
6-9
-------
drainage. In a few runs a problem was observed. When the
air flow rate is high and the liquid load is small, some
reentrainment is caused from the notch.
Liquid Catch Tanks
Four 0.2 m3 (55 gal) drums were used as catch tanks.
These were connected to the test section to collect sepa-
rated liquid alont the length of the test section. One
additional 0.2m3 drum was used to collect liquid from the
bottom of,the spray section. Each tank was connected to a
water meter and a pump with liquid level controller for re-
circulation of the liquid.
Liquid Supply Tanks
Three 0.2 m3 (55 gal) drums were used as supply tanks.
These tanks were connected to receive fresh water and re-
circulated water. On the outlet side were located rotameters
for flow measurement and pumps.
Control Panel for Equipment
The control panel was equipped with the following:
1. Electrical connections
,A. Magnetic starter for 88 m3/min blower
B. Switches for pumps, heater, sampling pump,
observation lights, etc.
C. Temperature recorder
2. Non-electrical connections
A. Rotameters and water meters
B. Dry gas meter
C. Pressure gauges
D. Needle valve, diaphragm valve, globe valves,
and gate valves
E. Manometers to measure pressure drop.
6-10
-------
Electrical Supply Panel
A 110 V, 3 phase, 100 amp/phase electrical supply panel
was installed near the equipment site,
Water Supply
The maximum fresh water supply was 50 Jt/min (13 gpm) .
Water supply to fine spray nozzles:
Piston pump - 1P741 (Grainger catalog)
Maximum pressure - 34 atm (500 Psi)
Flow rate at maximum pressure - 11.4 £/min (3 gpm)
Motor rpm - 1,725
Motor KW 1.1 (1.5 HP)
A rotameter with +_ 21 accuracy was used to measure flow
rate. A needle valve and auxiliary line near the pump were
used to control flow rate. A pressure gauge was used to
measure static pressure.
Water supply to coarse spray nozzles:
Centrifugal pump - Model 165U (Barnes Pump)
Maximum pressure - 3.4 atm (50 Psi)
Flow rate at maximum pressure - 120 A/min (31 gpm)
Motor rpm - 3,450
Motor KW 1.1 (1.5 HP)
A rotameter with +_ 2% accuracy was used to measure flow
rate. A diaphragm valve and auxiliary line near the pump
were used to control flow rate, A pressure gauge was used
to measure pressure.
TEST SECTION
Five different types of entrainment separators were
studied:
1. Mesh
2. Packed bed
3. Zigzag baffles
4. Cyclone
5. Tube bank
6-11
-------
Entrainment separators 1 through 4 were selected
because they are the most common separators used in in-
dustry. The fifth separator, the tube bank, was studied
to verify claims that it operates with a low pressure drop,
gives high efficiency, and has a high reentrainment
velocity.
The design of the test section was kept flexible to
facilitate any changes which may be necessary. The housings
used for the mesh, packed bed, zigzag baffles and tube bank
had the same dimensions, and thus it was a simple matter to
interchange the test sections.
The test sections were provided with a SO cm x 90 cm
plexiglass window in front and a 25 cm x 90 cm plexiglass
window at the top for observation. The dimensions of the
test section were 30.5 cm x 61 cm cross-section and 122 cm
in length. A false bottom was provided along the length of
the test section. The spacing of 1.9 cm height along the
length of the test section was loosely packed with Raschig
Rings and was divided into four parts to segregate for later
measurement, the collected liquid along the length of the
test section. A 5 cm diameter p.v.c. drainage pipe was pro-
vided in each section. A 30 cm liquid seal was provided in
each catch tank.
The five entrainment separators are described below.
Mesh
Model-4CA (ACS Industries)
Type - layered (with layers crimped in alternate
directions)
Density - 0.144 g/cm3
Wire diameter - 0.028 cm
Percent voids - 98.2
Mesh surface area - 2.8 cm2/cm3
Thickness - 10 cm
6-12
-------
Material of construction - AISI 304
Separation efficiency - Figure 6-5
Pressure drop data - Figure 6-6
The mesh was supported by a grid made from 0.16 cm
wire with 1,25 cm width and 2.5 cm height of spacing.
The mesh was located in the first 30 cm of the test
section.
Packed Bed
Packing - 2.5 cm pall rings
Specific surface - 1.9 cm2/cm3
Density - 0.088 g/cm3
Material of construction - Polypropylene plastic
Packing was supported by the same grid which is used
to support the mesh.
Some experiments were conducted using a 90 cm long bed.
Later the bed length was reduced to 30 cm. In both cases
the support grid at the upstream of the packing was located
at the beginning of the test section.
Zigzag Baffles
Baffle dimension - 7.5 cm width and 61 cm height
thickness 0.16 cm
No. of rows - 6
Spacing between rows - 2.5 cm
Angle between baffle and air flow direction - 30°
Spacing between baffles in a row - 6.9 cm
For details see Figure 6-7.
Cyclone
Diameter - 61 cm
Overall height - 244 cm
Inlet height - 30.5 cm
Inlet width - 15 cm
Maximum inlet velocity - 30 cm/sec., higher using an
inlet vane
6-13
-------
100
H
U
PH
w
w
H-l
1-5
o
0
2.5
GAS VELOCITY, m/sec
5.0
Figure 6-5 - Calculated collection efficiency for
water droplets in air. (ACS Industries)
6-14
-------
100
•
U
s
U
ex,
«
CO
CO
ex,
5 10 50
AIR VELOCITY, m/sec
100
Figure 6-6 - Pressure drop versus air velocity
for 10 cm thick ACS style 4CA mesh
L/A = superficial liquid velocity,
cm/min (ACS Industries)
6-15
-------
f i
in_L.
OM'i
\\\
////
v\ \ M
i / / /
CM
i
in
o
m
t f
30.5
air flow
Figure 6-7. Baffle Section. All
dimensions in cm. Aluminum baffles
are 0.16 cm thick, 7.5 cm wide, and
61 cm long.
6-16
-------
For details see Figure 6-8 and Stearman and Williams
(1971) .
Tube Bank
Number of rows - 6
External diameter - 1.9 cm
Length - 61 cm
Number of tubes in a row - 8
For details see Figure 6-9.
CALIBRATION
1. Thermocouples: All the thermocouples were calibrated
using thermometers. Ice-water, boiling water and room temp-
erature were used. It was found that thermocouples give some
errors in reading temperatures.
2. Liquid level in the tank: The liquid level as a
function of liquid height in the 0.2m3 drums was determined
by measuring the liquid level and the weight of the tank.
The liquid level in the catch tanks was controlled within
5 cm during the experiment. Tanks Tj, T3, T., T-, T,
contained 2.55 a/cm-of height, and tanks T2, A^, AT2
contained 5.06 £/cm of height (see Figure 6-3).
3. Water Meters: Water meter number 1 (on tank T ) was
calibrated using a rotameter and a calibrated 0.2 m3 drum.
Calibration was done at flow rate intervals of 3.79 £/min
(1 gpm).
Water meter Numbers 2 to 6 were connected to the outlets
of catch tanks Numbers 2 to 6. To calibrate any water meter
e.g., number 3, the outlet from water meter Number 1 was con-
nected to the inlet of tank 3. Tank 3 was filled with water
so that the liquid level was between an arbitrary upper and
lower limit. The pump was kept on to control the liquid level
in tank 3. The calibration was done by comparing the readings
6-17
-------
I
Figure 6-8 - Cyclone assembly. All
dimensions are in cm.
6-18
-------
in
Ki:
1*3. 8»f3. 8**3. »j*3. 8*3. 8*K3. W«3. 8^
Di O i Q. O Q ; O:OO i
-
8>f«3.8>f'3.8>r'3.8n*3.8»|-3.8*r0.8 *•
0.95
air flow
Figure 6^-9 - Test section with bank of tubes,
All dimensions in cm.
6-19
-------
on water meter numbers 1 and 3. Appropriate correction was
applied for the change in liquid level in tank number 3 and
for the calibration of water meter number 1. To avoid the
dynamic error, the calibration was done at flow rates of
3.79 Jl/min (1 gpm) intervals,
EXPERIMENTAL PROCEDURE
The major points of the experimental procedure are
described, below. The procedure was modified as required for
individual runs.
1. All the tanks were emptied before starting the
experiment in order to avoid rust in the water.
2. All the wet bulb thermometers were checked for
water.
3. All the valves were checked so that the required
valves were kept open and the rest closed.
4. All the recycle pumps were kept on to maintain the
liquid level in the tanks between the upper and
lower limits.
5. All the catch tanks were filled with liquid until
the level was between the upper and lower controlled
limits.
6. The feed supply tank was filled to the overflow
line. Fresh water was continuously added to the
system during the experiment to supplement the
water loss in the exit air and to introduce fresh
water to the system.
7. The zero position of the inclined manometer was
adjusted.
8. Readings were noted for all the water meters and
the liquid levels in the catch tanks.
9. The desired air flow was started.
6-20
-------
10. Pressure drop across the test section was
measured.
11. The desired water flow rate was started.
12. About 1-5 minutes were needed to reach steady
state. The experiment was continued for 2 hours.
13. Air flow rate, water flow rate, etc., were
checked every few minutes.
14. Visual observations of penetration, flooding,
liquid drainage, bouncing of drops, liquid flow
on elements of the entrainment separator, etc.,
were made for the duration of the experiment.
15. Readings were taken of temperature (each hour),
entrainment drop size, pressure drop, entrainment
loading, etc. (once during each run).
16. At the end of the experiment, liquid and air flow
were shut down. Readings were noted for water
levels in the tanks and water meter readings.
The blank forms used for the experiments are shown in
Figure 6-10.
SAMPLING PROCEDURE
Drop size determination: The drop diameter in the
experiments varied between 40 to 2,000 ym. Chemically treated
filter papers were used to determine the drop diameters. The
grain size of the chemicals normally limits the lower diameter
by this method to 5-10 ym. Filters coated with the following
chemicals were studied for drop size measurement.
Potassium permanganate
Lissamine green
Erio green
1% potassium ferricyanide and ferrous
ammonium sulfate.
6-21
-------
Entrainment Separation Project
Experiment #
Date
Test Section:
Air velocity
Water flow rate
Time on:
Types of nozzles
Water flow =
Thermocouples
Before
°F
After
°C
1
cm/sec
, Pitot tube -
ft/min
gal/min, Water pressure
, Time off:
psi
, Number of nozzles
SJ
\L
in
I
2
Total liquid level
After
Before
Net
cc
X
3
min
4
x 3.79xl03
5
i
i
2
6
3
cc
gal
7
4
8
cc
9^
^^
^^
^^^
**
5
—^
^^
„
^*t'
Water meter
After
Before
Net
x!03x
cc
Total
1
2
AP = Dry = +0.05 =
APxO.826 =
Abs . Pressure =
Dry
3
cm Wet =
cm,
Wet
4
+ 0.
OS «
5
o
,
^
„
cm
cm
cm Hg.
Drop diameter in inlet =
Geom. Standard deviation
Water overflowing
mass average
=
1-2
jjm
2-3
3
-4
Figure 6-10 - Form for recording test data.
6-22
-------
Reentrainment § Penetration
Approximate drop size of large drops mm
Vertical distance cm Horizontal distance cm
Height of flooded section cm
Reentrainment present in observation
Section yes no
Vertical height in which reentrainment
Present cm
No. of seconds sample taken
No. of drops on the filter paper
Isokinetic sample taken yes no
Drop diameter _ave. (mass)
Geometric Standard Deviation
COMMENTS:
Figure 6-10 - continued.
6-23
-------
Individual drops were most easily seen when 1% potassium
ferricyanide and ferrous ammonium sulfate were used as the
coating. A correlation given by Chilton (1952) Figure 6-11,
was used to convert blot diameter on filter paper to actual
drop diameter.
In a few runs isokinetic sampling through a filter holder
containing a treated filter was used to measure entrainment
drop diameter. This method gave good results when the entrain-
ment load was sufficiently small for individual drops to be
seen, or the gas velocity or drop diameter was such that the
collection efficiency of filter paper held in air is poor.
Determination of entrainment load:
1. Initially, the arrangement shown in Figure 6-12,was
used for liquid load measurement at the inlet and
outlet ends. Isokinetic sampling was used. It
was observed that all of the liquid was collected
in the inlet nozzle and none in the gravity sepa-
rator, cyclone or filter holder. The results
obtained were poor, which may be attributed to:
A. Entrainment may not be evenly distributed
in the cross-section of the duct.
B. Some liquid may have collected or drained
down from the sampling nozzle to the collection
flask during sampling.
2. The second method used to obtain entrainment loadings
was by material balance. The entrainment in the in-
let was obtained by subtracting liquid collected on
the walls before the test section from the liquid
feed and applying a correction for the humidity of
the ambient air. The entrainment in the outlet was
obtained by subtracting the liquid collected in
various parts of the test section from the inlet
entrainment.
6-24
-------
DROP DIAMETER, microns
M ^
o ' en "o "c
0 S ° C
0 0 0 C
'«—
x
/
;
^ ^
/
^
/
1
/*
-r*
/^
•
^
•
^
;;
0.3 O.S 1 5
BLOT DIAMETER, mm
10
Figure 6-11 - Calibration of Whatman No, 1
filter paper.
6-25
-------
Isokinetic Sampling
Air
Flow
gravity
separation
1
catch
collection
Filter Holder
xrptameter
cyclone
j| b needle/
/\valve "
LA
0
IT
dry
gas
'^ manometer
catch
collection vacuum pump
Figure 6-12 - Liquid load sampling
6-26
-------
This method was discarded due to errors in the
liquid balance caused by a fluctuation in the
inlet feed rate during the experiment. The
fluctuation may have been caused by the nozzles.
These errors may have been 5 to 10%. However,
in some experiments they may have been magnified
because liquid collected on the walls before the
test section was as high as 80% of the feed. Also,
some errors were found in the humidity correction
due to erroneous functioning of the thermocouples.
A third method was based on counting the number of
drops collected on a filter paper. The entrainment
load could be determined if the mass average drop
diameter and the sampling time were measured. This
method has the following disadvantages:
A. As shown in the previous chapter, the drop
collection efficiency by the filter paper varies
with the entrainment drop diameter and gas
velocity and is not 100%.
B. The error which is introduced in the measurement
of drop diameter is magnified when the volume
of the drop is calculated to determine the
entrainment load.
C. An error is introduced in inserting and removing
the filter paper from the duct because of the
short sampling time. When the entrainment load
is high, the sampling time is one second or less.
D. The sample represents entrainment load at a single
sampling point in the duct which may not represent
the average entrainment load.
6-27
-------
E, Some error is also introduced due to overlapping
of blots on the filter paper.
4. A method based on measuring the entrainment by con-
verting it to vapor and measuring the humidity of
the gas was selected for use in all experiments. An
isokinetic sample was drawn through a nozzle which
was heated to evaporate the entrainment. The air
sample humidity was measured by dry and wet bulb
thermometers following the nozzle. A thermostat
was used
-------
Transformer
Electrical
supply
Wet and dry
bulb thermometers
Impactor
•
Rotameter
Needle valve
Vacuum pump
Figure 6-13 - Sampling device consisting of impactor,
heated inlet probe, dry and wet bulb
thermometer and accessories
6-29
-------
was 3.8 Jl (1 gal.) per experiment, a maximum error of about
one half of one percent.
2. Liquid level. The maximum error in reading the
liquid level was about 2 mm, or one half of one percent. This
corresponds to 500 cm3 of liquid in catch tanks numbers 3 to 6
and 1 £ of liquid in catch tank number 2.
3. Water meter reading. All the water meters were
calibrated as a function of average flow rate (continuous
in water meter #1 and discrete flow in the rest of the water
meters). However, reproducibility was poor and the maximum
error was 5%.
4. Rotameter reading. The rotameter error was specified
as less than 2% of the maximum reading. However, additional
errors may have resulted during the experiments due to
fluctuation and drift in the flow. The two rotameters have
a maximum capacity of 8.4 £/min and 84 5,/min.
5. Temperature. The maximum error in temperature
reading with a thermometer was 1°C.
6. Air flow rate. Air flow rate was measured with a
pitot tube connected to an inclined manometer. The inclined
arm of the manometer was scaled to give velocity. Two types
of error were present: a) Error due to fluctuation in the
air flow rate, and b) Error due to change in fluid density
caused by a temperature change. The maximum error due to
temperature change was 50 cm/sec or 3-41, corresponding to
change of air velocity of 20 cm/sec in the test section.
6-30
-------
EXPERIMENTAL RESULTS AND DISCUSSION
At the time of writing, experiments had been run
on the pilot plant with horizontal gas flow through
packed bed, zigzag baffle, tube bank and mesh entrainment
separators. In each run air was used as the gas, water was
the liquid, and no solids were present. Results on overall
collection efficiency, pressure drop, reentrainment and
liquid flow patterns are presented and discussed in this
chapter. However, before the performance of various test
sections can be considered, the size distribution of the
inlet entrainment must be determined.
INLET ENTRAINMENT
Various nozzles were used in the experiments, although
only one type of nozzle was used in any given experiment.
A description of the nozzles is given in Table 7-1. A complex
relationship among the characteristics of the individual
spray nozzles, the interaction of multiple nozzles, the con-
figuration of the experimental duct and the air velocity
determines the inlet entrainment size distribution.
The spray generated from the M6 nozzles was analyzed
under experimental conditions by filter papers coated with
II potassium ferricyanide and ferrous ammonium sulfate, as
described in Chapter 5. The drop diameter generated from the
other nozzles was greater than 100 ym. For these, the manu-
facturer's data were used to determine drop diameters.
The effect of gas velocity on mass median drop diameter
generated from M6 nozzles is shown in Figure 7-1. There is no
definite trend. The mass median diameter varies from 76 to
102 ym and averages 84 ym, with an average geometric standard
7-1
-------
Table 7-1. NOZZLES USED IN SPRAY SECTION
Type of Nozzle
Hollow cone
Fogjet Nozzle
Full Cone
Hollow Cone
Model #
(Spray Systems)
1/4 M6SS
1-11 1/2 F18
1 1/2-11 1/2
F35
1/8 GG3
1/4 M26
i
Pressure,
atm gauge
13.6
2.7
2.7
2.7
2.7
# o£
Nozzles
12
1
1
12
12
&/m/nozzle
0.853
68
133
4
2
I
to
-------
100
w
H
o
oS
Q
C/3
0
10
GAS VELOCITY, m/sec
Figure 7-1 - The effect of gas velocity on
drop diameter for M6 nozzles
7-3
-------
deviation of 1.32. The minimum drop diameter found in the
inlet was 30 ym.
Inlet entrainment measurements were made at a point
75 cm downstream of the spray nozzles and 30 cm upstream of
the test section. The average water supply pressure was
13.6 atm. gauge (200 psig) . The nozzles were oriented toward
the downstream side and gave the drops an initial velocity
of 51 m/sec,
An analysis of the drop diameters created by each nozzle
is given in Table 7-2, and more detailed information concerning
the size distribution curves, as provided by the manufacturers,
is presented in Figures 7-2 through 7-4. These distributions
were measured 30 cm from the nozzles.
In these experiments the M26 nozzles were operated at
2,7 atm pressure, but the drop size data provided by the
manufacturer are for 6.8 atm and 10.2 atm. The mass median
drop diameter produced by M26 nozzles was obtained from
fitting the following relation for the effect of operating
pressure on drop diameter:
where d = mass median drop diameter, cm
mm
AP = pressure drop at nozzle, atm
c. , c2= constant
The mass median drop diameter for an operating pressure
of 2.7 atm was 380 ym. The geometric standard deviation was
1.52 and did not significantly vary with operating pressure.
The nozzles often plugged, due to formation of rust in
the water tanks. This resulted in a decreased water flow
rate and also may have caused some variation in the drop
diameter and standard deviation.
7-4
-------
Table 7-2. DROP SIZE ANALYSIS
I
en
Source of
Data
Manufacturer
Manufacturer
This Study
Manufacturer
Manufacture r
Predicted From
Equation (7-1)
Manufacturer
Type of
Nozzle
M6
M6
M6
M26
M26
M26
GG3
Operating
Pressure
atm gauge
6.8
10.2
13.6
6.8
10.2
2.7
2.7
Mass Median
Diameter,
ym
127
110
84
295
265
380
1,230
Minimum Drop
Diameter,
ym
45*
45*
30
110*
102*
-
450*
Geometric
Standard
Deviation
1.5
1.5
1.3
1.5
1.5
1.5
1.8
!2% of the drops are smaller than this diameter.
-------
w
H
O
erf
Q
1,000
500
300
200
100
50
30
20
10
T-I—i i i i i i i—rn—r
i i i i i i i i i i j i
2 5 10 20 50 80 90 95 98
ACCUMULATED VOLUME, %
Figure 7-2 - Drop diameter versus volume percentage for
hollow cone nozzle spraying water at 10.2
gauge pressure. (Manufacturer's data)
7-6
-------
E
a.
2
Q
a,
o
1,000
500
300
200
100
50
30
20
10
1 I I I I I I I I I I I I
II I I I I I I J I II
2 5 10 20 50 80 90 95 98
ACCUMULATED VOLUME, %
Figure 7^3 - Drop diameter versus volume percentage
for hollow cone nozzle spraying water at
6.8 atm gauge pressure, (Manufacturer's data)
7-7
-------
oo
<
)—I
Q
CU
§
3,000,
2,000
1,000
500
400
300
I I
J L
10
T—i—r
i—i—r
T—r
GG3
1 I i I »
J L
20 30
50
70 80
90
ACCUMULATED VOLUME, %
95 98
Figure 7-4 - Drop diameter versus volume percentage for
fulljet nozzles spraying water at 2.7 atm gauge
pressure. (Manufacturer's data)
-------
It was observed that the entrainment flow rate reaching
the entrainment separator decreased with decreasing air
velocity. This is due to an increase in collection by the
walls of the spray section.
EXPERIMENTAL RESULTS
Experimental data (see Appendix) for 153 runs are given
in Tables A-l to A-4. Table A-l gives the overall perfor-
mance of the entrainment separator. Liquid to gas volumetric
ratio and the entrainment mass median drop diameter are
measured at the inlet of the test section. Collection
efficiency is the overall efficiency and represents the
effect of primary efficiency, reentrainment and secondary
collection. Similarly, pressure drop is the overall pressure
drop for the test section.
Table A-2 gives the drop diameters of the entrainment
entering and leaving the test section. Since the water
feed rate, water pressure and type and number of nozzles
affect drop diameter in the inlet, these variables are in-
cluded in this table. The entry type and number of nozzles
give the nozzle model number followed in brackets by the
number of nozzles. The details of these nozzles are given
in Table 7-1.
Table A-3 gives the liquid material balance in the sys-
tem. As explained in the details of the equipment, the liquid
collected in the test section is divided into four parts
along the length of the test section. Liquid collected on
the side of the spray section and the observation section
upstream of the test section is also measured.
In some experiments liquid flowing as entrained drops
after the test section is determined by material balance,
i.e., [water in air after test section] = [total water
supplied] - [total water collected]. In the rest of the
experiments, liquid flowing as entrained drops was measured
7-9
-------
directly by measuring the humidity of a heated outlet gas
sample.
Table A-4 gives pressure drop data for the system.
"APd "'-and "APwetfl represent the pressure drops in the
test section at the experimental condition without and
with liquid flowing.
Packed Bed
Overall Efficiency - Efficiency data for packed beds are
presented in Figure 7-5. Runs at gas velocities (Note:
All gas velocities are superficial.) lower than 6.0 m/sec
did not show any penetration. The theory for primary col-
lection efficiency, shown as a solid line, is based on
Equation (3-13) and predicts 1001 primary efficiency over
the range of gas velocities studied. The data for overall
efficiency agrees well with the theory for primary efficiency,
which indicates that little or no reentrainment is taking
place.
Pressure Drop - Figures 7-6 and 7-7 show the pressure drop
in packed bed separators. There is no effect of liquid
load on pressure drop. The slope of the straight line on
the log-log graph is 1.94.
Reentrainment - According to Figure 7-8, there is no direct
dependence of the maximum outlet drop diameter on the gas
velocity in a packed bed. Figure 7-9 shows that the geo-
metric standard deviation increases with increasing outlet
mass median drop diameter. In both figures, the inlet mass
median drop diameter is used as a parameter, but no trends
can be detected as it is varied.
The minimum and maximum outlet drop diameters are
important in consideration of design criteria for a second
7-10
-------
100
ss
w
o
I— I
E-
Inlet Drop Diameter, ym
V 84
O
D
0
0
8
10
GAS VELOCITY, m/sec
Figure 7-5 Experimental collection efficiency of a
packed bed as a function of gas velocity
with horizontal flow.
7-11
-------
10
e
u
5 si
0.5
0.2
0.1
1.0 5 10
GAS VELOCITY, m/sec
Figure 7-6 - Dry pressure drop in packed
bed versus gas velocity
7-12
-------
10.0
5.0
§
0.2
0.1
1 5 10
GAS VELOCITY, m/sec
Figure 7-7 - Wet pressure drop in packed bed
versus gas velocity
7-13
-------
1,OQO
800
600
400
200
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Inlet Drop Diameter,
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0 1,230
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Figure 7-8 - Maximum drop diameter in the
entrainment versus gas velocity
for 30 cm packed bed
7-14
-------
1,000 m--±-j=jgJHg
Q
a,
CO
CO
w
100
10
Inlet Drop Diameter, ym
1.4 1.8 2.2
GEOMETRIC STANDARD DEVIATION
2.6
Figure 7-9 - Mass median drop diameter versus
geometric standard deviation
downstream of 30.5 cm packed bed
7-15
-------
stage entrainment separator. For packed beds the minimum
outlet drop diameter ranged from 40 to 60 ym.
Zigzag Baffles
Overall Efficiency - The overall collection efficiency for
horizontal gas flow through vertical zigzag baffles was
determined as a function of gas velocity. Figures 7-10
through 7-12 are plotted for different drop diameters in
the inlet entrainment. The separator attains 100% efficiency
for gas velocities between 3.0 and 6.0 m/sec. The efficiency
falls sharply for gas velocities below 3.0 m/sec for experi-
ments in which the inlet entrainment consisted of drops with
84 ym mass median diameter. Figure 7-13 shows data for all
the runs with zigzag baffles.
A theoretical curve for collection efficiency for
liquid droplets which have a median drop diameter of 90 ym
and geometric standard deviation of 1.35 is also shown in
Figure 7-10, where:
n = number of rows
9 = angle between the superficial and actual
air flow directions in the baffles
d = inlet mass median drop ciameter, cm
Pg
a = geometric standard deviation of inlet
o
entrainment
Experimental results reported by Bell and Strauss (1973)
for zigzag baffles are plotted in Figure 7-10 along with
points obtained in this study for d - 380 ym and a line
representing the data of Houghton and Radford (1938). The
inlet entrainment of the Bell and Strauss experiments was
comparable to this study, but their overall efficiency was
much lower. This is probably due to the differences in
separator design as reported in Table 7-3 .
7-16
-------
u
§
w
CJ
w
8
100
20
0
0123456
GAS VELOCITY, m/sec
Figure 7-10 - Collection efficiency versus gas velocity in the
zigzag baffle device with n = 6 and 9 = 30°.
The solid line represents theoretical curve for
d = 90 ym and a =1.35.
XT O O
7-17
-------
U
z
w
§
1-1
H
U
W
o
0
100
80
60
40
20
0
L£ Houghton § Radford
: Data, n=6, 6=30°
-*-i-t- f f
Strauss
tData for 2 "V"
^Baffles in
01 3456
AIR VELOCITY, ra/sec
8
Figure 7-11 -
Collection efficiency versus gas velocity in
the zigzag baffle device with n = 6, 9 = 30°,
= =
d = 380 ym and a = 1.52,
6
7-18
-------
U
I— I
PH
HH
w
-z
o
U
w
o
100
44iii:
80
60
40
20
0
M-i
L: t:.
F'Fit
_."4T "* rt4
T TO"
FF.
rtt+ ±111
•F
r,.ixiLL
t711
tr
.
tfe
Hi
F?:
-*iti
1^
F-ffi
$r
11
. -
(•rt
-
11
"T
. t, r j-4_._ _;
^FiM
tffi
fl
-:S
:; -rrtrt
1 3456
GAS VELOCITY, m/sec
8
Figure 7-12 - Collection efficiency versus gas velocity in
zigzag baffle device with n = 6t g = 30°,
d * 1,230 ym and a = 1.8. Theory predicts
1001 efficiency.
7-19
-------
100
80
w
I— I
E 60
o
M
H
O
CJ
40
20
~i~ T
< • • i
s
: ' j f
' 1 . /
• !/:.;-|
i •
" "!~ *!".".•"
1
. . _
i
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.
•" I "i !
[ • • l
• • 1
... . ,
.... - . . j
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- '
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- • i - 1 •
!
. i
- •
i
- r-
fc .
----- '-_'—'•
. _ i .
. . .-!--- -:
:-•• i : ;.
Inlet Drop Diameter, urn
:
' ; "" ' i
- -----
-
;
A
_ O
D
O
| '•''
'
... :
84
380
1,230
>1,230
I '.
i
!
i
t .
— : — •
4—4-'
' . .' :
' " : :.
, . . . . i .
i "
34567
GAS VELOCITY, m/sec
9 10
Figure 7-13 - Experimental collection efficiency as a
function of gas velocity in the vertical
baffles. Solid line represents theory.
7-20
-------
Table 7-3 .COMPARISON OF BAFFLE TYPE ENTRAINMENT SEPARATORS
Number of rows
9°
Lip to prevent
reentrainment
Staggering of rows
Distance between
rows
Spacing between
baffles in a row
Width of baffles
Present
Design
6
30
none
2 .5 cm
2.5 cm
6.9 cm
.7,5 cm
Bell $ Strauss
(1973)
4
45
1.9 cm on 1st
$ 3rd row only
none
3 .1 cm between
2nd § 3rd row
only
8.8 cm
6.2 cm
Houghton §
Radford (1938)
6
30
0.5 on 4th §
5th row only
none
0
2 cm
5 cm
7-21
-------
Houghton and Radford's experiments were conducted
under two operating conditions: (1) Liquid flow rate
=38 cm3/min and spray drop diameter ranging from about
1 to 60 ym, the predominant size being 40 ym, and
(2) Liquid flow rate = 12.3 fc/min and spray drop diameter
ranging from 2 to 800 ym, the predominant size being about
300 ym. The results obtained under both conditions were
similar and were comparable with the present results due
to similarities in the design, as summarized in Table 7.3.
Pressure Drop - Experimental dry and wet pressure drop
versus gas velocity for baffles is plotted in Figures 7-14
and 7-15, respectively. In both figures the solid lines
represent the theoretical prediction of pressure drop as
presented in Equation (4-8). The liquid load does not
have a significant effect on the pressure drop in the
baffle section. This should be expected as liquid hold-
up in the baffles is small because of the high drainage
rates.
Reentrainment - Figures 7-16 and 7-17 show the experi-
mentally determined dependence of the outlet mass median
drop diameter on the gas velocity. For an inlet drop
diameter of 280 ym, the outlet drop diameter increases
with rising gas velocity, as pictured in Figure 7-16.
However, Figure 7-17 reveals that no pattern is in evi-
dence for an inlet drop diameter of 1,230 ym.
The size distribution of drops leaving a zigzag baffl®
separator is depicted in Figure 7-18, with liquid flow rate
as the parameter. The two curves have nearly identical
slope on the log-probability diagram, but the lower liquid
flow rate resulted in larger reentrained drops.
The maximum drop diameter found in the outlet of a
baffle separator is graphed versus gas velocity for two
inlet drop diameters in Figure 7-19. A correlation can
7-22
-------
CJ
PL,
<
ttFn' iJ'' MMiiin i''iiif iliiiiiiiii'
ttrtttn ;i' m nr'il1 H i MUM! ill
02
5 10
GAS VELOCITY, m/sec
Figure 7-14 - Dry pressure drop in baffles
versus gas velocity
7-23
-------
5.0
u
U
2.0
.02
2345 10
GAS VELOCITY, m/sec
Figure 7-15 - Wet pressure drop in
baffles versus gas velocity
7-24
-------
1,200
w
H
§
o
CO
H
W
1,000
800
600
400
200
01 34
GAS VELOCITY,
5 6
m/sec
Figure 7-16 - Outlet drop diameter versus gas velocity
for zigzag baffles with inlet mass median
diameter of 380 ym.
7-25
-------
1,000
w
o
Pi
fi
en
W
O
800
600
400
200
I
-!I-l
*t|tr;
4:::
1 n
._-t-_;
M
5th
Km
IT;.
123456
GAS VELOCITY, m/sec
Figure 7-17 - Outlet drop diameter versus gas velocity
for zigzag baffle with 1,230 ym inlet
mass median drop diameter.
7-26
-------
1,000
500
6
.i
100
50
10
O Air velocity 6 m/sec, water flow rate 13.2 Jl/min
Air velocity 6 m/sec, water flow rate 7.6
i
m
•!!'
•
U-4-
5
T 4-1
UJTuT
i i • • i ;
10 20 30 40 50 60 70 80 90
CUMULATIVE NUMBER, %
f
95
98
Figure 7-18 . size distribution of drops leaving baffle entrainment
separators
7-27
-------
i— i
tm
O
o
O
pm
Pi
w
H
UJ
PH
o
Pi
Q
i—)
X
H
W
J
H
D
O
00
o
O
ON
o
O
4*.
o
O
^j
o
CD
o
iF
-^:
t]
:-•)•'
mm
%&j&
i r
i •
T"
14332
-mr-
- ii'^rf •-:
k-+-
__|.
•: !
T^r~frr\~~.'
i'liiE.-lii -
T ~t~ i. i •
.-.7;-;;
^ Q
^J4^x±-
Inlet Drop Diameter, ym
380
1,230
I
-T-:-
; 1 • -
I j •" -"•
0
2
6
8
GAS VELOCITY, m/sec
Figure 7-19 -
Maximum outlet drop diameter in the
entrainment versus gas velocity for
zigzag baffles.
7-28
-------
be established for 380 ym drops but not for the large
1,230 ym drops. The minimum drop size in the reentrain-
ment was observed to range from 50 to 80 ym.
The geometric standard deviation gets larger as the
outlet mass median increases, as seen in Figure 7-20.
Two very important parameters for determining re-
entrainment are gas velocity and the liquid to gas ratio.
The combination of these two which results in reentrain-
ment as observed experimentally is shown in the shaded
region in Figure 7-21. The reentrainment observed is
0.5 - 1% of the inlet entrainment.
Liquid Flow on the Baffles - Flow of liquid on the baffle
surface was observed in some of the runs. As the flow
increases, the film thickness of the liquid near the down-
stream edge increases. The gas stream forms a wake at the
back side of the baffle, which tears away some of the liquid
at the downstream edge. The approximate shape of the wake
is shown in Figure 7-22. The wake formation becomes more
pronounced with increasing gas velocity. The flow of liquid
film on the back side of the baffle is shown in the same
figure.
If the liquid flow on the baffle surface is small, only
drop flow takes place on the back side of the baffle. Some
of these drops reach the upstream edge of the baffle, where
they are reentrained. The reentrained drops splash on the
adjacent baffle in the same row and disintegrate. Some of
these small drops are reentrained in the air. The drops
normally splash on the third quarter width of the baffles
as measured from the upstream. The drops flowing on the
back side of the baffles are 3-4 mm in diameter.
Reentrainment from the downstream edge of the baffle
was more significant compared to reentrainment from the
7-29
-------
10
Q
I
Q
Q
W
CO
cn
H
W
o
10
10
. I . -1 . ;
H : -f . '"^- r •:-••
1.4 1,8 2.2
GEOMETRIC STANDARD DEVIATION
2.6
Figure 7-20 - Drop diameter versus geometric standard
deviation at zigzag baffle outlet.
7-30
-------
10
- 3
6
\
CO
6
10
en
o
H
10
- 5
A Some reentrainment (<1%)
i . .
XN Reentrainment in part of I . ;":
_J_. N/duct only ;
Primary efficiency <100%
T f i ' ~ j~
: O No penetration
: i > l f' i ; t - ; r' I • •—-•••-- i ; . = : i- : : ! ; "TT
3456
GAS VELOCITY, m/sec
Figure 7-21 Effect of gas velocity and liquid
load on performance of baffle type
separator.
7-31
-------
Formation of wake
Wall acting as collector
.1
Pulsating
liquid flow
on the back
side of the
baffle
Figure 7-22 - Some observed phenomena in entrainment
separator (a) formation of wake
(b) liquid flow on the back side of
the baffle (c) wall effect
7-32
-------
upstream edge. If the liquid flow on the baffle surface
was drop flow, some of these drops reached the downstream
edge and (1) were reentrained, (2) were turned to the
back side of the baffle, (3) fell down at the edge due to
gravity, or (4) stayed at the edge of the baffle until
they grew by coalescing with other drops. Most of the
drops were collected by the third or fourth step. If the
liquid was flowing as a film on the baffle, part of the
film was torn and reentrained at the downstream end. The
drops reentrained from the downstream edge of the baffle
were 3-5 mm in diameter. These drops were normally col-
lected on the baffles of the second row, i.e., drops re-
entrained from the second row were collected on the baffles
in the fourth row.
Some wall effect was observed in the baffle section.
There were four baffles in a row and the side walls of the
test section acted as collectors for the entrainment. This
effect is shown in Figure 7-22.
The liquid flow pulsated whenever reentrainment took
place and occurred in film flow and in drop flow. It was
difficult to determine the amplitude of the pulsating film
which may have been of the order of 0.05 cm. The frequency
of the wave was not measured.
Mesh
Overall Efficiency - The overall collection efficiency data
for horizontal flow through wire mesh are plotted in
Figure 7-23. No penetration was observed in experiments
at low gas velocity, less than 3.0 m/sec. At higher velo-
cities penetration due to reentrainment was observed. The
dotted line, predicting 1001 efficiency, represents the
theoretical curve based on Equation (3-21). In the range
of experimental data, the curve given by the manufacturer
7-33
-------
i nn
on
oU
vs
u
Jz;
M
U
HH
£ 60
w
COLLECTION
•t*
o
20
0
(
1 | | II! ! | I
w-J V V-Jy VI "'Bjty \/ "i11 iy Ljj V ^pv 1aiHjytj'"\/
';• /' • • • I ;• • ; ; ^ i
i : . ' I .! j
:':: •': -'•'•:: j ;: • \ : ;
; . . : • '
— • • ' , :• ...... .. ._
! ' ' ; ' ' j
Inlet Drop Diameter, urn
V 84
O 380
E 1,230 i
O > 1,230 ;
,11111111
J.I 23 4 56788 9 1C
GAS VELOCITY, m/sec
Figure 7-23
Experimental collection efficiency of
wire mesh as a function of gas velocity.
7-34
-------
(ACS Industries) for mesh efficiency corresponds to the
theoretical curve.
Pressure Drop - In contrast to the results for packed beds
and zigzag baffles, the pressure drop in wire mesh sepa-
rators is affected by liquid load, as seen in Figure 7-24.
The slope of the straight lines on the log-log plot is 1.65;
thus, "AP" can be represented as a function of "lu1'6511. In
Figure 7-24, L/A = 0 represents the dry pressure drop
"APd " through the mesh. For 0 < L/A < 1 the pressure
drop is 1.5 AP, , and for 1 < L/A < 5 the pressure drop
is 2.3 APdry.
Reentrainment - The outlet mass median drop diameter is
plotted as a function of gas velocity in both Figures 7-25
and 7-26. Each curve shows a dependence in the range of
gas velocities studied, but it is more pronounced for
inlet drop diameters of 82 ym than for large drops of
1,230 ym diameter.
Figure 7-27 demonstrates that no trend can be detected
relating gas velocity with the maximum drop diameter in the
reentrainment. However, the minimum drop diameter was found
to range from 40 to 80 ym. As seen in Figure 7-28, there is
a positive straight-line correlation between geometric
standard deviation and the mass median drop diameter in the
outlet.
Figure 7-29 shows the combination of gas velocity and
liquid to gas ratio which will result in reentrainment in
mesh separators. The shaded area is the region where
onset of reentrainment was observed.
Buerkholz (1970) collected reentrainment data for
sulfuric acid mist 150 cm downstream of a mesh separator.
He found that reentrainment increased from 1.6 to 4.0$
(0.3 to 1.3 mg/m3) as the gas velocity was increased from
7-35
-------
3.0
u
0
U
1 < L/A < 5
lffliii!ill,!ilM:l'):;|;;:;ltJij£si:
0 < L/A < 1
GAS VELOCITY, m/sec
Figure 7-24 -
Pressure drop in wire mesh
versus gas velocity with liquid
load as parameter
L/A = Superficial liquid velocity
cm/min
7-36
-------
1,000
Pi
W
H
P
O
g
CO
LO
H
w
-J
H
§
800
600
400
200 -
0
12345
GAS VELOCITY, m/sec
Figure 7-25 - Outlet drop diameter versus gas velocity
for mesh with 82 pm inlet drop diameter.
7-37
-------
500
w
H
3
§
n
Q
w
2
CO
3
H
O
400
300
200 —
100 —-
2345 67
GAS VELOCITY, m/sec
Figure 7-26 - Outlet drop diameter versus gas velocity
for mesh with 1,230 ym inlet drop
diameter
7-38
-------
1,000
OH
O
X
,-J
H
800
600
400
200
0
0 3456
GAS VELOCITY, m/sec
Figure 7-27 - Maximum outlet drop diameter in the
entrainment versus gas velocity for
mesh
7-39
-------
ym
ETER
DROP
Q
W
S
co
H
W
P
O
M
0
10
1.4 1.8 2.2
GEOMETRIC STANDARD DEVIATION
2.6
igure7-28 _ Drop diameter versus geometric standard
deviation for mesh.
7-40
-------
1x10
- 3
e
CO
6
O
M
H
g
o
ex
HH
-J
1x10
IxlO"5
Reentrainment
Figure 7-29 -
GAS VELOCITY, m/sec
Effect of sas velocity and li
load on performance of mesh
7-41
-------
4.7 to 8.2 m/sec. The outlet mass median drop diameter also
increased from 150 to 750 ym. Buerkholz1 data, plotted in
Figure 7-30, were collected on a 15 cm x 15 cm mesh with
sedimentation present between the mesh and the sampling point.
The solid line in Figure 7-30 is the onset of reentrainment
curve obtained in the present study. The data show good
agreement in determining the reentrainment velocity of 5 m/sec
at very small liquid loads.
The reentrainment curve obtained from the manufacturer
also appears in Figure 7-30. The manufacturer predicts
higher reentrainment velocity than the present results. The
differences may be due to the fact that the manufacturer used
vertical gas flow.
Visual Observation of Reentrainment - Reentrainment in the
mesh section was observed to take place in the following ways:
1. At low liquid loads the mesh operated without
flooding. Apparently, the drops that are collected on the
mesh wires grew to 3-5 mm diameter before they drained down.
If the air velocity is high, the path of the drop is not
vertically downward. Some of these drops were airborne and
struck the wires of the grid supporting the mesh. Normally,
the drop shattered into one large drop slightly smaller than
the original size and 2-4 satellite drops which were
Some of the drops collected on the grid wire drained at
once, whereas the rest drained after growing to a larger size.
There were other drops which missed striking any wires and
emerged from the mesh. These drops were collected at the
downstream side of the mesh. The drops passing through the
mesh without striking any wires were carried farther down-
stream of the mesh than others. All these reentrained drops
were 4-5 mm in diameter and upon reaching the bottom, they
shattered into a few (3-4) satellite drops. The rest of the
liquid in the original drop was mixed with the liquid film
7-42
-------
10
- 3
6
M
o
I—I
E-
10
D
cc
10
0
8
GAS VELOCITY, m/sec
Figure 7-30
Onset of reentrainment velocity curves for
mesh. (—) Experimental data, (--) manu-
facturer's catalog, (0) Reentrainment data
observed by Buerkholz (1970),
7-43
-------
at the bottom. These satellite drops flew into the air due
to kinetic energy, and their initial trajectory formed a
cone along a vertical axis. The angle of the cone was de-
pendent upon initial drop velocity and was observed to range
from 0° to 90°. Some of these satellite drops were re-
entrained while others fell down.
2. Some drops were reentrained inside the mesh, and
the process of reentrainment could not be observed; it is
assumed to be the same as described in the earlier part of
the first method.
3. When the liquid load was high, partial flooding was
observed. Reentrainment by methods 1 and 2 took place above
the flooded zone. In the flooded section the air flow rate
was low. The flooded section was partially covered by the
falling drops from above on the downstream side of the mesh.
The reentrainment mechanism was rupture of bubbles, but it
could not be observed properly. Reentrainment by the above
mechanisms was observed in runs 43 and 46.
Tube Bank
Overall Efficiency - Collection efficiency versus gas velocity
data for horizontal flow through tube banks are plotted in
Figures 7-31 through 7-33 for various inlet drop diameters.
Penetration due to primary efficiency of less than 100% was
observed for velocities lower than 3.0 m/sec.
Houghton and Radford's (1938) data for strut separators
are also plotted in Figure 7-32. They found a constant col-
lection efficiency of 96.2% for gas velocities from 1.25 to
17.5 m/sec. The inlet entrainment contained drops as small
as 1 ym. However, no increase in penetration at lower velo-
cities or reentrainment at higher velocities was observed. A
comparison between the configuration used in the present
study and that of Houghton and Radford is given in Table 7-4'
7-44
-------
o
hH
H
U
w
o
u
0
8
GAS VELOCITY, m/sec
Figure 7-31 - Collection efficiency versus gas velocity
in tube bank with n = 6, d = 84 urn and
a = 1.32.
7-45
-------
100
Houghton $ Radford Data (1938)
|: OWater loading 17.1 £/min
M Awater loading 6.8 £/min
loading 11,4
OWater loading 3.8 £/min
o
123456
GAS VELOCITY, m/sec
Figure 7-32 - Collection efficiency versus gas velocity in
tube bank with d = 380 ym and a =1.5
7-46
-------
100
w
PU
H
U
W
O
u
80
60
40
20
^jIWWI Hi .' ' :::
^£?M | h
-T;-}.^---» -:- r-;-~
:, :i^_ " i .' :
fZ7 ST ;
i
-:| .:. : : 1
- | • ' 1,230
. 1 :'- i . ' . . ; ..
1 . j ,j.,,.,
M^npisi
^^_i^^^ ,^. ' jf JT^T1^ .^^ • ; 1 '
/^ ' X\
^5 >i^
1 i • • •
m. ' " . ' ' . j ' . '. '
• . '..,._; 1 I
< ; •• . i-:':|/:;
meter, ym
, . . ,4.1.. . j. . - , . t, .-.- i —
J
• • ; " «" : •
i < ' ' •
1
' 1 - ' ' ' ~
1- ! 1 1
234567
GAS VELOCITY, m/sec
Figure 7-33 - Collection efficiency versus gas
velocity in tube bank. Solid line
represents theory.
7-47
-------
Table 7-4. COMPARISON OF TUBE BANKS
Present
Study
Houghton §
Radford (1938)
Tube (minor-major
axis) diameter, cm
Spacing between
tubes, center to
center, cm
Spacing between
rows, center to
center, cm
Number of rows
Material of
tubes
1.9
3.8
3.3
6
Al
1.25x3.2
2.5
7-48
-------
Pressure^ Drop - Dry and wet pressure drop through the tube
bank is plotted in Figures 7-34 and 7-35. There is no
effect of liquid load on the pressure drop; the slope of
both curves is 2.33. Therefore the data may be correlated
by the following equation:
AP = l.Z.xlO-5 n pQ vG2'33 (7-2)
where n = number of rows (6 in present experiments)
v^ = actual gas velocity near the tubes
The actual gas velocity "vV! is twice the superficial
gas velocity, which is shown on the horizontal axes of
Figures 7-34 and 7-35. An average value of 1.17x10"3 g/cm3
is used for MPG"«
Manufacturer's data (Union Carbide Corporation for
M.V. separator) for banks of streamlined struts of 2,2 cm
equivalent diameter gave the following empirical equation
for pressure drop:
AP = 2.33x10-6 n PG Vg2'19 (7-3)
A comparison of Equations (7-2) and (7-3) shows that the
former equation gives 5-8 times higher pressure drops for
cylindrical tubes with 1.8 cm equivalent diameter than for
streamlined struts with larger equivalent diameter.
Grimison (1937) has presented friction factor as a
function of tube bank configuration at various Reynolds
numbers. For the configuration studied in the pilot plant,
friction factor versus superficial gas velocity is plotted
in Figure 7-36. Substituting the value of friction factor
into the pressure drop relation results in:
AP - 4.33x10-* n PG Vg1'9 (7-4)
Equation (7-4) predicts pressure drop twice as high as that
observed in the present study for VG = 6.0 m/sec.
7-49
-------
2.0
CJ
*
3=
6
A
<
.02
.01
1 5 10
GAS VELOCITY, m/sec
Figure 7-34 - Dry pressure drop in tube
bank versus gas velocity
7-50
-------
2.0
u
o
PL,
<\
.02
.01
GAS VELOCITY, m/sec
Figure 7-35 -
Wet pressure drop in tube
bank versus gas velocity
7-51
-------
0.15
M-i
1.0 2345 10
GAS VELOCITY, m/sec
50
Figure 7-36 - Friction factor versus gas velocity
in the bank of tubes
7-52
-------
Reentrainmeiit
Figures 7-37 and 7-38 reveal no definite relationship
between gas velocity and outlet mass median drop diameter.
However, the size distribution curves shown in Figure 7-39
are similar to those obtained from zigzag baffles in
Figure 7-18. For both types of separators, a higher liquid
flow rate caused smaller drops to be formed.
Figure 7-40 gives a similar relationship between geo-
metric standard deviation and outlet mass median drop
diameter as seen for each of the other three types of .
separators studied. The maximum outlet drop diameter
rises with increasing gas velocity, according to Figure 7-41,
to a peak of about 350 ym at 7.0 m/sec gas velocity. The
minimum drop diameter at the outlet, on the other hand,
ranged between 40 and 70 ym.
Figure 7-42 depicts the values of gas velocity and
liquid load observed as being necessary for reentrainment.
As in the case of zigzag baffles (Figure 7-21), the re-
entrainment region has an upper and lower gas velocity
boundary for a given liquid load, whereas for the mesh
(Figure 7-29) there is only a lower boundary.
GENERAL OBSERVATIONS,
Pressure Drop
The pressure drop in the empty test section housing as
a function of gas velocity is plotted in Figure 7-43. The
housing for each separator studied thus far has the same
dimension. When the pressure drop is less than 1 mm W.C.,
it is essentially negligible. Therefore, there is no sig-
nificant pressure drop for a gas velocity less than about
5.0 m/sec.
7-53
-------
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^.tff.
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-*• —
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H::::
.._ ..
".^r :.
——
0
4567
GAS VELOCITY, m/sec
Figure 7-37 -
Outlet drop diameter versus gas velocity
for tube bank with 380 ym inlet drop
diameter.
7-54
-------
OS
w
H
O
P^
O
&
0
120
en
E-
UJ
-J
H
3
O
Bip Nozzles Used ±g
±ttt
123456
GAS VELOCITY, m/sec
'igure 7-38 - Outlet drop diameter versus gas velocity
for tube bank with 1,230 ym inlet drop
diameter.
7-55
-------
1,000
500
100
on
w
H
Q
O
50
10
.•-:t££p O Air velocity 6 m/sec, water flow rate 19.0 £/min
Air velocity 6 m/sec, water flow rate 11.4 £/min
--!i
.1. »-* --1 ..-F-i
i i i i IIiIi
:i
J
rn
i !
..:_L.
10
30 40 50 60 70
CUMULATIVE NUMBER, %
80
90
95
98
Figure 7- 39 - Size distribution of drops leaving tube bank
entrainment separators.
7-56
-------
10
w
H
W
OH
O
Di
P
t—i
Q
CO
CO
UTLET
M
O
10
Inlet Drop Diameter, ym
A 84
O
El
380
1,230
1 1.4 1.8 2,2
GEOMETRIC STANDARD DEVIATION
2.6
Figure 7-40
-Outlet mass median drop diameter versus
geometric standard deviation for tube
bank.
7-57
-------
1,000
n
w
O
(X,
1
s
o
800
600
400
200
0
- • • —
iFifcc
ffffaf
I.JH-Hli:
Inlet Drop Diameter, ym
A 84
0
8
GAS VELOCITY, m/sec
Figure 7-41 - Maximum drop diameter in the entrainment
versus gas velocity for tube bank.
7-58
-------
1x10
6
•v.
€n
6
•k
O
i—i
H
CO
o
H
Q
CX
i—i
*-l
1x10-"
1x10
A ;Some reentrainment
n part duct
E£
O No penetration
or <$> penetration
due to less than
100% primary
efficiency
GAS VELOCITY, m/sec
Figure 7-42 Experimental results showing the
effect of gas velocity and liquid
load on performance of tube bank.
7-59
-------
u
0.1
0.05
GAS VELOCITY, m/sec
:igure 7,43 - Pressure drop in the empty
section versus gas velocity
7-60
-------
During the operation of the entrainment separator, the
velocity distribution changes due to the presence of the
separator. Thus, the wall effects measured by the above
method may not represent the true wall effect. However,
the presence of the separator does reduce the wall effect.
Also, the pressure drop in the empty test section is 10-20%
of the pressure drop with entrainment separator. Hence,
the effect of the walls on the total pressure drop is small.
The dry pressure drop versus gas velocity relationships
for different entrainment separators are shown in Figure 7-44
The packed bed is seen to have the greatest pressure drop.
It should be noted that the efficiencies of different sepa-
rators at a given velocity are not the same.
Bell and Strauss' pressure drop data for baffles are
compared with present results for baffles and tubes in
Figure 7-45. The high pressure drop found by Bell and
Strauss can be explained by the variations in design, dif-
ferences in "0" and the distance between rows and the use
of lips.
The pressure drop data of Houghton and Radford are
compared with the present results in Figure 7-46. The
results show good agreement. The small differences in the
pressure drops may be explained as below.
The differences in the design are given in Tables 7-3,
7-4, and 7-5. The higher pressure drop in the baffle sec-
tion obtained by Houghton and Radford may be due to lips
on the fourth and fifth row of baffles and no spacing
between the rows.
Houghton and Radford's pressure drop data in the tube
section give comparable results at velocities lower than
500 cm/sec. At higher velocities the present study gives
higher pressure drops, which may be due to smaller spacing
between rows and to the use of tubes with circular cross-
section rather than struts with streamlined cross-section.
7-61
-------
8
5
S
u
0.5
0.3
0.2
0.
0.05
0.03
1 5 10
GAS VELOCITY, m/sec
Figure 7-44 - Dry pressure drop versus gas
velocity used in pilot plant
7-62
-------
12
w
o
H
PH
U4
CO
*
CJ
6
u
10
w
OS
CO
8
0
23456
GAS VELOCITY, m/sec
Figure 7-45 - Pressure drop through entrainment
separator as a function of superficial
gas velocity
7-63
-------
u
•
12
•V
PH
=Zigzag Baffles
Mesh
0.1
5 10
GAS VELOCITY, m/sec
Figure 7-46 - Comparison of pressure drop data of
Houghton § Radford (1938) with present
results. The dashed lines represent
the present results. The three curves
for the mesh represent L/A = 0,
0 <_ L/A <_ 1, 1 < L/A 1 5. Here, "L/A"
is superficial liquid velocity, cm/min.
7-64
-------
Table 7-5. COMPARISON OF WIRE MESH DEMISTER
Mesh Density,
g/cm3
Void space
Wire diameter,
cm
Bed length,
cm
Material
Miscellaneous
Present
Results
98.2
0.028
10
SS
Bell § Strauss
(1973)
0.16
0.025
10
SS
Houghton §
Radford (1938)
Number 4 mesh
layers kept
5 cm apart
7-65
-------
Reentrainment
Entrainment in the outlet is due to less than 100%
primary efficiency or reentrainment. In some runs at low
gas velocity, 1.2-2.4 m/sec, entrainment occurred because
primary efficiency was less than 100%. Increasing the gas
velocity caused the entrainment to disappear in the outlet.
Entrainment due to low primary efficiency occupied the
entire cross-section of the duct and was observed in runs
5, 146, 147, and 152.
The mass median drop diameter of the reentrainment
varied between 70 and 1,000 ym with a geometric standard
deviation of 1.2 to 2.4. Some drops settled by gravity
between the test section and the position downstream where
drop diameters were measured. These measurements were
taken at the estimated middle of the cross-section in which
entrainment was present, resulting in a variation of the
sampling point elevation. No correction for these two
effects has been made in the analysis of drop diameters.
As noted earlier in this chapter, the mass median drop
diameter increases with increase in geometric standard
deviation. Thus the drop diameter distribution becomes
broader as the median diameter increases. Similar results
were obtained by Garner et al. (1954). They attribute large
values of geometric standard deviation to reentrainment
caused by shattering of drops. The resulting drops are
larger than 200 urn with mass median diameter between 250 and
1,000 ym.
The effect of liquid feed rate on the typical outlet
mass median drop diameter and on the reentrainment velocity
for different separators is summarized in Table 7-6. The
typical diameter and reentrainment velocity are both seen
to rise as the liquid feed rate is decreased. Figures 7-38
and 7-39 confirm the data of Table 7-6 for baffles and tube
7-66
-------
Table 7-6. EFFECT OF LIQUID LOAD ON REENTRAINMENT IN DIFFERENT
ENTRAINMENT SEPARATORS.
Test Section
Tube Bank
Tube Bank
Mesh
Mesh
Zigzag Baffles
Zigzag Baffles
Typical Outlet
Mass Median
Diameter
90 vim
350 ym
170 ym
260 ym
700 ym
700 ym
Average Liquid
Feed
cm3/sec
4xl02
2.7xl02
4xl02
l.SxlO2
4xl02
2.7xl02
Reentrain-
ment "Velo-
city m/sec
1.5
3.5
1.8
4.5
1.5
3.0
7-67
-------
banks, respectively. In the case of mesh, experimental data
show that the mass median drop diameter increases with liquid
load.
The minimum diameter of reentrainment for each type of
separator is given in Table 7-7, which is a distillation of
the data plotted in Figure 7-47. In a few experiments with
baffles, the minimum size was as high as 190 ym. As shown
in Figure 7-47, most data points lie between 40 and 80 ym.
Kotov (1972) has summarized reentrainment from various
separators (Table 7-8) . Reentrainment velocities in the
present study were much higher than those obtained by Kotov.
However, the experimental details of his work are not avail-
able, and it is difficult to make comparisons.
7-68
-------
Table 7-7. OBSERVED MINIMUM DROP SIZE
IN THE REENTRAINMENT
Separator
Baffles
Mesh
Packed Bed
Tube Bank
Minimum
50
40
40
40
Drop Diameter
urn
- 80
- 80
- 60
- 70
7-69
-------
200
*
rt
H
CM
O
OS
n
HH
2
S
L '
s
H->
en
0
M
O
0
en
0
o
GAS
4567
VELOCITY, m/sec
Figure 7-47 - Minimum outlet drop diameter versus
gas velocity
7-70
-------
Table 7-8. REENTRAINMENT OBSERVED BY KOTOV
Device or
Mechanism
Liquid Load
In Outlet
m3/m3
Gas
Velocity
m/sec
Screen or Mesh
Ceramic Rings
Bubble Break Up
0.4 - O.SxlO6
0.83 - l.lSxlO6
0.7 - 8xl06
14 - 19xl06
<2.8
2.8 - 3.0
7-71
-------
7-72
-------
FUTURE RESEARCH AND DEVELOPMENT RECOMMENDATION
The primary objectives of the present research were to
evaluate the available technology on wet scrubber entrainment
separators, advance theoretical development and design
a pilot scale entrainment separator based on the above
information. It is also important to define the areas
where additional work is needed. At this time, the present
contract is in progress and the recommendations made in
the following paragraphs are subject to modification as
more is learned during the course of the research program.
REENTRAINMENT
One of the problems which present day entrainment
separators suffer is their large size which is due to low
operating velocities. The gas velocities are limited by
reentrainment velocities and flooding conditions. Reentrain-
ment may take place due to various mechanisms, depending
on flow rates and geometry.
While the present program will provide information on
the conditions under which reentrainment occurs in several
separator configurations, it would be helpful to have more
detailed knowledge of this phenomenon. It is quite
possible that a fundamental study of the mechanisms of
reentrainment from different geometric arrangements would
enable one to develop more efficient separator designs. At
least the results of such a study would delineate the limits
of performance possible and save effort which might otherwise
be expended in unprofitable directions.
The study needed is onset of reentrainment conditions,
rate of reentrainment, equilibrium constant between en-
trainment and liquid in film, drop size distribution, smooth
and shock type contact of gas and liquid, effect of duct
8-1
-------
dimensions, etc. The application to entrainment separator
will include improving design methods to determine re-
entrainment under operating conditions, effect of higher
gas velocities and improvements in design to reduce
reentrainment,
SOLID DEPOSITION
Solids deposition and consequent plugging is a major
operational problem in scrubber systems. It would be very
helpful to have insight into the mechanism(s) by which sus-
pended and dissolved solids deposit and the effects of
operating and design parameters. Given a clearer picture
of how solids deposit, it should be easier to conceive of
and design for conditions which would minimize deposition.
It is probably unavoidable that some deposition will
take place in a high efficiency entrainment separator.
Therefore, means for removing deposits are required. Washing,
either intermittently or continuously, is the most common
method used and it has the undesirable feature of introducing
liquid where it is not wanted. Research on methods of
washing and flow rate required would be worth doing, with
the objective of finding the optimum way to use the least
liquid.
FIELD TESTING OF INDUSTRIAL ENTRAINMENT SEPARATORS
Performance data on industrial entrainment separators
are generally not available. The industrial data are col-
lected to evaluate the overall performance of the scrubber
and it is assumed that the entrainment separators have 100%
efficiency. Also, all the liquid introduced in the wet
scrubber is assumed to be removed by entrainment separator.
The effects of sedimentation, bends in the duct carrying
entrainment, etc. are neglected. The distance between
8-2
-------
sampling point and entrainment separator elements is
important. Also, the effects of industrial operating
conditions on performance of entrainment separators should
be determined.
The aim of development of entrainment separators is
to improve performance of separators under industrial
conditions. Thus, it is necessary to collect data on
industrial separators. The data, when compared with
theoretical models, will represent possible problems re-
sulting from industrial conditions and will help in de-
signing future entrainment separators,
DEMONSTRATION PLANT
From the present contract work, it is felt that we
can determine the performance of an entrainment separator
with reasonable accuracy. It is possible to obtain im-
provement in the performance due to better design. We
would like to move from the present research and develop-
ment to a demonstration of an improved design in the field.
The capacity of the present pilot plant is 85 m3/min.
Therefore, the next size should be around 1,000 m3/min
(35,000 CFM).
The demonstration plant operation will involve
selecting an organization which operates a suitable plant
having entrainment separation problems and which is willing
to participate in the demonstration plant program. The
design effort will include obtaining the necessary data
concerning the source of entrainment, preparing overall
design and selecting a final design. The fabrication and
start up will involve selection and negotiation with sub-
contractors, procurement of components and supervision of
subcontractor efforts. The test program will be to deter-
mine performance, observe the effect of change in variables
and compare the performance with theoretical developments.
8-3
-------
TO STUDY COMBINATIONS OF ENTRAINMENT SEPARATORS
It seems that if more than one entrainment separator
is used, the combined series unit will offer a synergistic
effect. It is possible to combine two different entrain-
ment separators to include the best features of each. Some
examples are as follows:
1. The maximum gas velocity in the entrainment
separator is limited to the onset of reentrainment velocity.
It is generally the case, however, that a separator which
has high primary drop collection efficiency will have a low
reentrainment velocity, while one with high reentrainment
velocity will have low primary collection efficiency. If
a combined unit is used with the first unit being used for
primary collection and coalescence of drops and the second
one for collection of large drops while being below the
onset of reentrainment velocity, increased capacity will
result. The combination of efficiency and capacity will
exceed what either unit can do alone.
Because the size of the entrainment separator will be
smaller, the initial capital cost will be lower. The mini-
mum drop size that can be separated in the entrainment
separator is limited by the operating velocity. This prob-
lem can be solved by using a combination of entrainment •
separators.
2. Sometimes the entrainment load is high and con-
stituted of particles in a wide size range. A single
entrainment separator may be inefficient, flooded or may
present reentrainment in this situation. A combined unit
may be used in this case. The first separator is a pre-
cleaner with low pressure drop, which removes large particles
constituting a significant fraction of the entrainment. The
second separator will be an efficient device.
8-4
-------
APPENDIX
-------
Table A-l. OVERALL PERFORMANCE OF ENTRAINMENT SEPARATOR
Exp.
No.
1
2
3
4
5
6
8
9
11
12
13
14
15
17
18
19
20
21
22
23
Test
Section
Baffle
Baffle
Baffle
Baffle
Jaffle
'Baffle
Baffle
Tubes
Tubes
Tubes
Tubes
Tubes
Tubes
Tubes
Tubes
Tubes
Mesh
Mesh
Mesh
Mesh
Air
Velocity
m/sec
4.8
3.2
6.1
2.4
1.8
1.2
4.9
5.4
4.2
3.0
2.0
2.4
1.2
6.3
6.4
6.2
6.1
4.4
3.0
211
.4
Liquid to
Gas Ratio
m3/m3
3.4xlO-5
2.97xlO-5
5.65xlO-5
4.66xlO"s
4.95xlO"5
9.14xlO-'5
1.213xlOr"
1.133x10-"
N.A.
3.224xlO-5
N.A.
5.609xlO-5
1.122x10""
1. 676x10'"
N.A.
3.06x10-"
1.36x10-"
1.83x10-"
2.1x10'"
2.5x10-"
Median
Drop
Diameter
ym
97
94
97
102
93.6
87
79
76 /
84
83
82
78
1225
1225
1225
1225
1225
1225
1225
Collection
Efficiency
%
*100
*100
*100
*100
N.A.
69.2
=aoo
*100
*100
*100
N.A.
78.8
78.6
=aoo
*100
100
94
N.A.
>99.9
>99.9
Pressure
Drop
cm W.C»
1,7
0.41
3.3
0.41
.0.12
0.12
1.5
0.82
0.29
0.16
0.08
0.12
0
1.3
1.4
1.5
2.2 -
1.3
0.62
0.54
Reentrainment
Present in
Observation
Section
—
— '
fm
-
fm
^
mm
-
—
_
fm
-
_
_
^
-------
Table A-l. OVERALL PERFORMANCE OF ENTRAPMENT SEPARATOR
(continued)
•
Exp.
No.
24
25
26
27
28
29
30
31
32
33
34
35
36
Test J
Section 1
Mesh
Mesh
Mesh
Mesh
Mesh
Mesh
Tubes
Tubes
Tubes
Tubes
Tubes
Tubes
Tubes
37 (Tubes
38 [Tubes
40 Mesh
41 Mesh
43 Mesh
45 Mesh
46
i
i
Mesh
Mr
Velocity
m/sec
1.8
2.4
. 2
1.2
2.4
3.2
6.0
6.3
6.5
4.5
3.0
1.5
2.2
3.0
6.9
4.5
1.5
7.1
3.0
6.9
Liquid to
Gas Ratio
2.72x10-"
2.37x10-"
3.87x10-"
7.46x10'"
4.55x10-"
3.57x10-"
2.53x10'"
3.78x10-"
2.936x10-"
4.026x10-"
5.224x10-"
1.733xlO-3
5.661x10'"
1.297x10""
6.636xlO"5
2.579xlO-5
1.403xlQ-5
9.48x10- 6
5.685xlO-5
2.577xlO-5
edian
rop
iameter
ym
1225
1225
1225
1225
1225
1225
1225
1225
1225
1225
1225
1225
1225
380
380
84
84
. 84
84
84
Collection
Efficiency
i
>99.9
N.A.
. 100
100
90.3
87.8
N.A.
91
87.8
90.0
94.4
98.2
99.8
100
98.7
94.3
100
97.7
50
91.8
Pressure
Drop
cm W.C.
0.21
0.45
0.21
.0.29
0.54
0.78
1.6
1.8
1.3
N.A.
0.37
0.12
0.12
0_ *
.54
1.03
0.620
0.084
1.36
0.29
Reentrainment
Present in
Observation
Section
-
*•
—
—
-
*~
~
Some
"
Yes
••
~
—
—
-
—
Yes
*™
-------
Table A-l
OVERALL PERFORMANCE OF ENTRAINMENT SEPARATOR
(continued)
Bxp.
No.
47
48
49
50
51
52
53
54
55
56
58
59
60
61
63
64
65
66
67
69
Test
Section
Mesh
Mesh
Mesh
Tubes
Tubes
Tubes
Tubes
Tubes
Tubes
Tubes
Tubes
Tubes
Tubes
Tubes
Mesh
Mesh
Mesh
Mesh
Mesh
Mesh
Air
Velocity
m/sec
4.8
6.0
5.0
6.0
5.4
4.8
4.0
7.0
4.2
5.2
6.8
4.2
5.4
7.1
7.4
6.0
7.8
6.0
4.5
6.0
Liquid to
Gas Ratio
m3/m3
4.21xlO-5
3.532X10'5
2.448xlO-5
1.317x10-*
N.A.
9.782xl(T5
1.0127x10-"
9.374xlO-5
8.445xlO-s
8.442xlO-5
4.859xlO-5
5.739xlO-s
4.552xlO"5
2,171xlO-5
8.38X10-11
8.23X10-11
2, 34x10- 5
3.442xlO-5
4.616xlO-5
8. 354x10 -5
Median
Drop
Diameter
yra
84
84
84
380
380
380
380
380
380
380
380
380
380
380
84
84
84
84
84
84
Collection
Efficiency
%
64.8
63.1
80.8
100
• N.A.
100
100
83.2
93.7
95.1
100
94.94
98.6
74.6
61.5
57.6
100
73.22
100
Pressure
Drop
cm w.C,
0.87
1.3
0.87
0.87
0.87
0.66
0.55
1.5
0. 5
0.79
1.3
0.54
0.91
W • mJ J_
1.36
1.6
1.1
w <3
Yes
Some
Very little
Some
Some
Yes
* \* «j
Yes
Some
U V/1ILW
Yes
JL w O
Some
-------
Table A-l. OVERALL PERFORMANCE OF ENTRAINMENT SEPARATOR
(continued)
Exp.
No.
70
71
73
74
75
76
77
78
79
80
Test
Section
Mesh
36" Pall
36" Pall
36" Pall
36" Pall
36" Pall
Baffle
Baffle
Baffle
Baffle
81 Baffle
82 Baffle
83 Baffle
84 Baffle
85 Baffle
86 Baffle
88 baffle
89
90
91
Baffle
Baffle
Jaffle
Air
Velocity
m/sec
4.5
5.0
4.9
1.0
5.0
4.9
6.9
6.0
5.0
6.0
6.8
3.0
3.0
7.0
6.0
1.5
7.0
6.0
4.5
3.0
Liquid to
Gas Ratio
m3/m3
5.383xlO"s
6.607xlO-5
2.950x10-"
4.588x10-"
1.131xlO-3
1.553xlO-3
6.873xlO-5
4.733xlO"5
7.005xlO-5
5.849xlO-5
6.297xlO-5
7.850xlO-s
4.351x10-"
2.330x10-"
2.393x10-"
5.384x10""
1.602x10-"
1.718x10-"
2.784x10'"
3.501x10'"
Median
)rop
Diameter
ym
380
380
380
1,225
1,225
_
— »
380
380
380
380
380
380
1,225
1,225
1,225
1,225
1,225
1,225
1,225
Collection
Efficiency
%
100
100
100
100
90.8
100
80
97.7
100
100
77.9
100
100
100
100
100
100
100
100
100
Pressure
Drop
cm W.C,
0.54
8.9
9.3
0.37
4.2
9.5
3.2
2.5
1.8
2.4
2.9
0.5
0.5
3.0
2.2
0.45
2.9
2.0
1.2
0.45
Reentrainment
Present in
Observation
Section
Slight
No
No
No
No
No
Yes
Some
Very little
-
Yes
Little
Little
Some
Some
Little
Little
Little
Some
-------
Table A-l. OVERALL PERFORMANCE OF ENTRAINMENT SEPARATOR
(continued)
Exp.
No.
92
93
94
95
97
98
99
100
101
102
103
104
105
106
107
108
109
110
111
112
Test
Section
Baffle
Baffle
Baffle
12" Pall
12" Pall
12" Pall
L2" Pall
L2" Pall
I1 Pall
Rings
it
it
jt
i*
tt
it
it
tt
tt
tt
tt
Air
Velocity
m/sec
2.0
1.0
7.1
5.8
2.0
1.0
5.9
4.5
2.0
1.0
1.0
6.0
3.5
6.0
3.5
6.0
6.0
4.5
6.0
4.5
Liquid to
Gas Ratio
mVm3
4.281x10-*.
6.9i3xio-*
8. 008x10-*
9.672x10-*
2.038xlO'3
3.233xlO-3
1.310xlO-3
l,689xlO-3
2.48xlO-3
3. 86x10- 3
6.56x10- 5
2.28xlO-5
I. 33x10- 5
3.16x10- 5
1.37x10-"
2.76xlO-s
1.15x10"*
5.21xlO"5
2.88x10'*
3.53x10-*
Median
Drop
Diameter
ym
1,225
1,225
1,225
-
-
84
84
84
84
84
84
380
380
1,225
1,225
Collection
Efficiency
"1
100
100
100
100
100
100
100
100
100
100
100
88.7
100
100
100
100
93.7
100
N.A.
100
Pressure
Drop
cm W.C.
0 25
0 33
2.6
N.A.
0.54
0.12
7.6
% n
3 . U
0.45
0.17
0.21
6.5
2.2
6.5
2.3
6.6
6.5
3.7
6.6
4.2
Reentrainment
Present in
Observation
Section
Yes
Slight
No
No
No
No
No
Slight
No
No
No
No
Some
No
Some
No
-------
Table A-l. OVERALL PERFORMANCE OF ENTRAINMENT SEPARATOR
(continued)
o\
Exp. Test
No. Section '
113 1' Pall
Rings
m"
i
115 "
116
117
118 " •
119 "
i
120 "
•^ i
121
122
123 tesh
124
125
126
127
128
130
131
132
133
11
it
ii
H
it
rt
ti
it
it
kir
Velocity
m/sec
6.0
6.0
6.0
6.0
4.5
6.0
3.5
6.0
3.5
6.0
7.2
4.5
1.0
3.0
7.0
6.0
3.5
2.0
3.0
7.2
Liquid to 1
Gas Ratio
mVm3
2.85x10-"
2.20x10-*
2. 07x10-*
2.60x10'*
3.47x10"*
2. 5 8x10 -5
4.l2xlO"5
ll.67xlO-s
1.74x10-*
2.75xlO-5
8.03xlO-5
4.06xlO-5
4.67xlO"5
4.55xlO-5
2.48xlO'5
S.OOxlO"5
5.87xlO'5
8.81xlO-5
1.920x10-*
2.794x10'*
1
Median
Drop
Diameter
um
1,225
1,225
1,225
1,225
1,225
84
84
84
84
84
84
84
84
84
84
84
84
84
1,225
1,225
Collection
Efficiency
%
N.A.
N.A.
N.A.
N.A.
100
100
100
100
100
100
88.7
100
100
100
87.6
79.1
100
100
100
100
Pressure
Drop
cm W . C .
6.7
6.6
6.7
7.1
4.3
6.4
2.60
7*T
.3
2.5
6y-
.5
1.9
0. 87
0.04
0.29
1.9
1.4
0.45
0.12
0.54
1.9
Reentrainment
Present in
Observation
Section
Some
Some
Some
-------
Table A-l. OVERALL PERFORMANCE OF ENTRAINMENT SEPARATOR
(continued)
Exp.
No.
134
135
136
137
138
140
141
142
143
145
146
147
148
149
150
151
152
153
Test
Section
Mesh
it
11
"
it
it
it
it
it
11
Tubes
it
it
it
it
Baffle
it
Packed
Bed 12"
Air
Velocity
m/sec
5.0
6.6
6.8
6.8
6.8
2.4
3.2
6.0
5.0
7.0
1.2
2.4
7.0
6.4
4.5
6.8
1.2
6.0
Liquid to
Gas Ratio
mVm3
3.390x10-*
2.861x10"*
2.73x10-*
8.31x10-*
8.34xlO~*
5.62xlO~5
2.23x10'*
4.606xlO"5
5 .'836x10 ~5
4.96xlO-5
1.006x10'*
8.634xlO-5
1.099xlO"5
3.022x10-*
3.471xlO-5
6.&07xlO-5
6.627xlO-5
4.277xlO"5
Median
)rop
Diameter
ym
1,225
1,225
1,225
_
-
380
1,225
84
84
84
84
84
380
1,225
1,225
380
84
84
Collection
Efficiency
100
100
100
100
100
100
100
100
100
99
99.5
99.2
99.5
100
99.8
99.8
99.7
100
Pressure
Drop
cm W . C .
1.5
2.5
2.5
2.9
2.8
0.27
0.62
1.5
0.95
1.6
0.04
0.21
1.5
1.20
0.79
2.8
0.08
5.4
Reentrainment
Present in
Observation
Section
-------
Table A-2. DROP DIAMETERS OF THE ENTRAINMENT ENTERING AND
LEAVING THE TEST SECTION
Exp.
1
Water
Feed Rate
cm3/sec
9.48X101
2 J9, 48x10 J
3
4
5
6
8
9
11
12
13
14
15
"i *i
17
18
19
20
21
22
23
1.52x10*
1.52xl02
1.52xl02
1.52xl02
3.16xl02
1.77xl02
1.64xl02
1.48xl02
L.26xl02
L.33xl02
L.33xl02
1 ,21x10
2.21xl02
l.llxlO2
1.90xl02
L.90xl02
L;90xl02
L.90xl02
Water
Pressure
atm. ("gauge")
5.44.
5.78
13.61
14,29
14.29
14.97
3.40
N.A.
N.A,
N.A.
N.A."
N.A.
N,A,
3.33
1.95
3.27
3.27
3.33
3.33
3.33
Type § No .
of nozzles
M6SS(12)
M6SS(12)
M6SS(12)
M6SS(12)
M6SS(12)
M6SSC12)
M26C12)
M6SS(12)
M6SS(12)
M6SS(12)
M6SS(12)
M6SS(12)
M6SS(12)
GG3(6)
GG3(6)
GG3(12)
GG3(6)
GG3C6)
GG3C6)
GG3C6)
Inlet
Median.
Drop Dia-
meter
urn
97
94
97
102
.- 94
87
79
. 76
78
84
83
82
78
1225
1225
1225
1225
1225
1225'
1225
Geometric
Standard
Devi at ion
1.4
1.3
1.3
1.4
1.3
1.3
1.3
1.3
1.3
1.3
1.3
1.3
1.3
1.8
1.8
1.8
1.8
1.8
1.8
1.8
Outlet 1
Median Drop
Diameter
N.A.
N.A.
-
93
86
70
-
91
120
. 91
Geometric
Standard
P ej/ iation
N.A.
N.A.
_
1.3
1.3
1 7
JL . £•
..
1.3
1.2
1.3
-------
Table A-2. DROP DIAMETERS OF THE ENTRAPMENT ENTERING AND
. . . LEAVING THE TEST SECTION (continued)
Ixp.
Jo.
t
24
25
26
27
28
29
30
31
32
33
34
35
36
37
38
40
41 .
43
45
46
Water
Feed Rate
cmVsec
1.90xl02
M
1. 90x10*
1.90xl02
3.16xl02
3.16xl02
3.16xl02
f*
3.16xl02
4.11xl02
3.45xl02
4.00xl02
4.01xl02
3.98xl02
4.01xl02
2.15xl02
1.94xl02
6.64xlOl
•
5.36X101 .
4. 75x10 l
M
l.OlOxlO2
1.072xl02
Water
Pressure
atm. ("gauge")
3.33
3,33
3.33
3,33
3.33
3.33
3.33
3.33
3.40
3.40
3.4
3.4 '
3,4
3.4
3.4
13.95
13.6
13.6
13.6
13.6
Type g No,
o£ nozzle;
GG3(6)
GG3(6)
GG3(6)
GG3(9)
GG3(9)
GG3(9)
GG3(9)
GG3(12)
GG3(12)
GG3(12)
GG3(12)
GG3(12)
GG3(12)
M26(12)
M26(12)
M6SS(6)
M6SS(6)
M6SS(6)
M6SS(12)
M6SS.(12)
Inlet
Median
Drop Dia-
meter
urn
1225
1225
1225
1225
1225
122*5
1225
1225
1225
1225
1225
1225
1225
380
380
84
..84
84
84
84
Geometric
Standard
Deviation
1.8
1.8
1.8
1.8
1.8
1.8
1.8
1.8
1.8
1.8
1.8
1.8
1.8
1.5
1.5
1.3
1.3
1.3
1.3
1.3
Outlet i
Median Drop
Diameter
}im
91
* ^
_
96
76
83
115
94
125
109
97 . '
103
_
330
87.6
232
211
Geometric
Standard
Devi nt ion
1.3
1.3
1. 2
1.3
1.4
1.3
1.4
1.4
1,3
1.3
1,6
1.2
1.4
1.7
-------
Table A-2.
DROP DIAMETERS OF THE ENTRAINMENT ENTERING AND
LEAVING THE TEST SECTION (continued)
iixp.
*0.
47
48
49
50
51
52
53
54
55
56
58
59
60
61
63
64
65
66
67
68
Water
Feed Rate
cm3/sec
1.074xl02
9.736X101
7.264X101
3.056xl02
2.833xl02
2.61xl02
2.41xl02
1.914xl02
1.714xl02
1.78xl02
1.267xl02
1.161xl02
1.069xl02
5.472x10*
3.79X101
3.79X101
l.lOSxlO2
l.lOSxlO2
1.074xl02
2.595xl02
i
Water
Pressure
atm, f gauge")
13.6
13.6
13.6
3.40
3.40
3.40
3.40
3.40
3.40
3.40
3.40
3.40
3.40
3.40
13.6
13.6
13.6
13.6
13.6
3.4
Type § No,
of nozzles
M6SS(12)
M6SS(12)
M6SS(12)
M26(12)
M26(12)
M26(12)
M26(12)
M26(9)
M26(9)
M26(9)
M26(6)
M26(6)
M26(6)
M26(3)
M6SS(3)
M6SS(3)
M6SS(9)
M6SS(9)
M6SS(9)
M26(12)
Inlet
Median
Drop Dia-
meter
urn
84
84
84
380
380
*
380
380
380
380
380
380
380
380
380
84
84
84
84
84
380
Geometric
Standard
Deviation
1.3
1.3
1.3
1.5
1.5
1.5
1.5
1.5
1.5
1.5
1.5
1.5
1.5
1.5
1.3
1.3
1.3
1.3
1.3
1.3
Outlet
Median Drop
Diameter
jim
1377
123
414
317 '
413
1063
300
300
336
392
87
78
308
109
264
389
252
-
109
m
Geometric
Standard
Deviation
2.3
1.4
1.7
1.7
1.8
2.3
1.6
1.6
1.6
1.8
1.3
1.2
1.8
1. 5
1.7
1.9
1.8
1.4
-------
Table A-2.
DROP DIAMETERS OF THE ENTRAINMENT ENTERING AND
LEAVING THE TEST SECTION (continued)
ixp .
4o.
69
70
71
73
74
75
76
77
78
79
80
81
82
83
84
85
86
88
89
90
Water
Feed Rate
cm3/sec
2.333xl02
2.198xl02
2.164x10*
3.881xl02
3.564xl02
1.079xl03
1.459xl03
1.165xl02
1.272xl02
2,027xl02
1.763xl02
1.767xl02
1.761xl02
3.644xl02
3.594xl02
3.439xl02
3.442xl02
A
2.661X102
A
2.567xl02
3.194xl02
.
Water
Pressure
atm. f gauge")
3.40
3.40
3.40
3.40
3.40
Type $ No,
of nozzlei
M26(12)
M26(12)
M26(12)
1/8"GG3(12
1/8"GG3(12
3.06 [small gard
fiose (1)
2.04 [large gard
hose (1)
3.40 M26(12)
3.40 M26(9)
3.40 M26(12)
3.40
3.40
3.40
3.40
3.40
3.40
3.40
3.40
3.40
3.40
426(12)
^26(12)
H26(12)
1/8"GG3(12
1/8"GG3(12
1/8"GG3(12
1/8"GG3(12
1/8"GG3(9)
1/8"GG3(9)
1/8"GG3(9)
Inlet
Median
Drop Dia-
meter
urn
380
380
380
) 1,225
) 1,225
sn
sn
380
380
380
380
380
380
) 1,225
) 1,225
) 1,225
) 1,225
1,225
1,225
1,225
Geometric
Standard
Deviation
1.5
1.5
1.5
1.8
1.8
_
-
1.5
1.5
1.5.
1.5
1.5
1.5
1.8
1..8
1.8
1.8
1.8
1.8
1.8
Outlet
Median Drop
Diameter
urn
115
—
«.
-
—
—
1080
798
688
419
634
162
714
3025
1060
981
614
454
511
Geometric
Standard
Devi at ion
1.4
—
_
-
.
_
2 2
£* • £*
1 8
J- • LJ
2.1
1.7
1.6
1.4
1.6
2.4
1.9
1.6
1.7
1.7
1.7
-------
Table A-2. DROP DIAMETERS OF THE ENTRAINMENT ENTERING AND
LEAVING THE TEST SECTION (continued)
Exp.
91
92
93
94
95
97
98
99
100
Water
Feed Rate
cmVsec
3.206xl02
3.192xl02
3.15xl02
1.086xl03
1.069xl03
1.077xl03
1.078xl03
1.454xl03
1.453xl03
Water
Pressure
atm. ("gauge")
3.40
3.40
3.40
3.12
3.20
3.20
3.20
2.04 1
2.04
Type $ No
of nozzlei
L/8"GG3(9)
L/8"GG3(9)
L/8"GG3(9)
small gard
lose (1)
>mall gard
lose (1)
small gardi
lose (1)
>mall gardi
lose (1)
,arge gardi
lose (1)
,arge gardi
lose (1)
'
Inlet
Median
Drop Dia-
meter
urn
1,225
1,225
1,225
m
*'
in
;n
;n
Geometric
Standard
Deviation
1.8
1.8
1.8
Outlet
Median Drop
Diameter
um
795
696
N.A.
N.A.
95
88
^_
Geometric
Standard
Deviation
1.8
1. 5
N.A.
N.A.
1.3
1. 3
-------
Table A-2. DROP DIAMETERS OF THE ENTRAINMENT ENTERING AND
LEAVING THE TEST SECTION (continued)
C/4
Bxp.
to.
101
102
103
104
105
106
107
108
109
110
111
112
113
114
115
116
117
118
119
120
Water
Feed Rate
cm3/sec
1.453xl03
1.449xl03
1.263xl02
1.137xl02
1.137xl02
7.264X101
9. 159x10 l
3.474X101
1.327xl02
3.097xl02
4.453xl02
4.492xl02
4.233xl02
3.050xl02
3.058xl02
3.567xl02
3.878xl02
1.137xl02
1.137xl02
7.264X101
Water
Pressure
atm. (gauge 1
2.041
2.041
13.605
13.605
13.605
13.605
13.605
13.605
3.401
3.401
3.401
3.401
3.401
3.401
3.401.
3.40
3.40
13.61
13.61
13.61
Type § No.
of nozzle*
11-1/2F35
11-1/2F35
M6(12)
M6(12)
M6(12)
M6(6)
M6(6)
M6(3)
M26(12)
M26(12)
GG3(12)
GG3(12)
GG3(12)
GG3(9)
GG3(9)
GG3(12)
GG3(12)
M6(12)
M6(12)
M6(6)
Inlet
Median
Drop Dia-
meter
urn
-
84
84
-
84
84
84
380
380
12
12
12
12
12
12
12
84
84
84
Geometric
Standard
Deviation
-
1.3
1.3
-
1.3
1.3
1.3
1.5
1.5
1.8
1.8
1,8
1.8
1.8
_
-
1.3
1.3
1.3
!
Outlet
Median Drop
Diameter
ym
-
-
69
-
_
-
294
510
-
298
-
.
281
351
277
-
-
-
Geometric
Standard
Deviation
-
-
1. 3
-
_
-
1.6
1.8
-
1.6
-
1.5
1.6
1.6
-
-
-
-------
Table A-2. DROP DIAMETERS OF THE ENTRAINMENT ENTERING AND
LEAVING THE TEST SECTION (continued)
Exp.
to.
121
122
123
124
125
126
127
128
130
131
132
133
134
135
136
t
137
138
140
141
142
Water
Feed Rate
cm3/sec
1.137xl02
3.474X101
1.074xl02
1.074xl02
1.074xl02
1.074X102-
1.074xl02
1.074xl02
1.074xl02
1.074xl02
3.494x10*
4.536xl02
4.392xl02
4.261xl02
4.182xl02
1.067xl03
1.069xl03
2.S65xl02
4.185xl02
1.263xl02
,
Water
Pressure
atm. (gauge)
13.61
13.61
13.61
13.61
13.61
13.61-
13.61
13.61
13.61
13.61
3.40
3.40
3.40
3.40
3.40
3.06
3.06
3.40
3.40
13.60
Type § No.
of nozzles
M6(12)
M6(3)
M6(12)
M6(12)
M6(12)
M6(12)
M6(9)
M6(9)
M6(9)
M6(9)
1/8GG3(12
1/8GG3(12
1/8GG3(12
1/8GG3(12
GG3(12)
11-1/2F18
11-1/2F18
M26(12)
GG3(12)
M6(12)
Inlet
Median
Drop Dia-
meter
urn
84
84
84
84
84
84
84
84
84
84
1 1225
I 1225
I 1225
) 1225
1225
-
-
380
1225
84
Geometric
Standard
Deviation
1.3
1.3
1.3
1.3
1.3
1.3
1.3
1.3
1.3
1.3
1.8
1.8
1.8
1.8
1.8
-
-
1.5
1.7
1.3
Outlet
Median Drop
Diameter
ym
_
-
798
-
-
-
605
152
-
-
-
-
-
441
472
-
-
-
-
"
t
Geometric
Standard
Deviation
_
-
1.9
-
-
-
1.9
1.5
-
—
-
-
-
1.6
1.6
-
-
-
-
-------
Table A-2. DROP DIAMETERS OF THE ENTRAINMENT ENTERING AND
LEAVING THE TEST SECTION (continued)
Exp.
4o.
143
145
146
147
148
149
150
151
152
153
Water
Feed Rate
cm3/sec
I,263xl02
1.390xl02
1.273xl02
1.334xl02
2.353xl02
4.321xl02
4.214xl02
2.118xl02
1.314X102
I,300xl02
Water
Pressure
atm. f gauge ")
13.60
13.60
13.60
13.60
3.40
3.40 "
3.40
3.40
13.60
13.60
Type § No
of nozzles
M6(12)
M6(12)
M6(12)
M6(12)
M26(12)
GG3(12)
GG3(12)
M26(12)
M6(12)
M6(12)
Inlet
Median
Drop Dia-
meter
urn
84
84
84
84
380
1225
1225
380
84
84
Geometric
Standard
Deviation
1.3
1.3
1.3
1.3
1.5
1.8
1.8
1.5
1.3
1.3
Outlet
Median Drop
Diameter
urn
86
87
-
Geometric
Standard
Deviation
1.3
1.3
-
I
M
tn
-------
Table A-3. LIQUID MATERIAL BALANCE
Exp.
No.
1
2
3
4
5
6
-8
9
.12
13
14
15
17
18
19
20
21
22
23
24
Input
cm 3
6.07xl05a
3.329xlOsa
7.61xlOsa
5.36xl05
7.91xl05
7.28xl05
1.367xl06a
1.337xl06
5.076xl05
N.A.
3.741xl05
4.775xl05
5.63xl05
N.A.
L.232xl06
5.12X105
i.481x!05
7.618X105
?.391xl05
5.912xl05
\
i
Before
Test Section
% of Input '
68.2 •
81.3
57.7
84.4
93.9
91.1
64.9
35.6
88.1
3.226xl05
85,0
85.7
7.8
4.42xl05
13.9
18.7
21.1
38.1
41.1
52.0
Test Section
Section I
•6 of Input
17.1
14.2
27.0
10.1
5.8
5.2
18.9
64, Oa
10.9
3.659x10*
10.9
5.4
28.3
5.438xl05
63.3
23.9
49.7
60.4
61.7
65.0
Section .11
* of Input
12.9
0.4
7.4
2,7
0.6
- 0 . 6
8.0
0.3
12.1
3.94x10*
11.9
5.8
16.5
3.31Sxl03
0
23.5
6,7
0.4
0
0
Section III
% of Input
1,7$
. 0
5.46
0
0
0.4
7.33
0
0
0
0
0
19.78
2.805xl033
0
25.93
2.1
0
0
0
Section IV
% of Input
0.043
0
1.63
0.053
0
0
0.83
0
0
0
0
0
0.23
3.545x10*3
0
3.13
1.9
0
0
0
After Test
Section in air
?o of Input
0
0
0
0
N.A, '
2.7a
0
0
0
N.A.
3. 2ct
• *•* w»
3 . 5ct
0
N.A.
0
4.9
N.A.
1.1
N.A.
N.A.
(See notes on page A-23)
-------
Table A-3. LIQUID MATERIAL BALANCE .(continued)
Exp.
No.
25
26
27
28
29
30
31
32
.33
34
35
36
37
38
40
41
43
45
46
47
Input
cm3
6.822xl05
5.685xl05
1.137xl06
L.137xl06
1.137xl06
4. 738x10 5
1.3S5xl06
5.917xlOs
L.470xl06
L.481xl06
2.888xl06
L.470xl06
7.954xl05
7.915xl05
2.285xlOs
1.895xl05
1.545xl05
3.568xl05
3. 70 5x10 5
3.754xl05
i
\
I
Before
Test Section
% of Input '
44.3 •
54.4
47.3
35.7
32,8
10.6
41.2
7.6
17.5
26.8
39.4
41.3
67.2
47.5
66.0
92.6
71.0
68.0
67.9
64,0
-
Test Section
Section I
1 of Input
60.3
56.8
50.8
57.4
34.1
31.5
27,4
36.1
62.5
63.5
-47.7
28.9
26.3
21.0
32.0
9.1
15.7
16.0
10.2
20.6
Section J3
% of Input
0
0
0
0.4
16,8
' 29.23
6.63
453 '
18.63
5.53
11.33
27.53
12.3
17.2
0
0
12.8
0
19.33
2.6
section III
% of Input
0
0
0
0.33
9.53
14.83
11.13
0
0.5
0
0.1
0
0
0
0
0
0
0
0
0.06
Section IV
% o£ Inrmt
0 -
0
0
0
6.93
19.23
8. '33
0
0
0.2
0
0.1 '.
0
0
0
0
0
0
0
0
After Test
Section in air
% of Input
N.A.
0
0
6.2
14.9
N.A.
5,3
11.3
0.9
4.1
1.5
2.2
0
r.92a
0
0.5a
16. Oa
2 ,6a
12. 7a
(see notes on page A-23)
-------
Table A-3. LIQUID MATERIAL BALANCE (continued)
Exp.
No.
48
49
50
51
53
54
55
56
57
58
59
60
61
63
64
65
66
67
68
69
Input
cm 3
3.365xl05
2.499xl05
1.126xl06
1.007xl06
8.567xl05
6.727xl05
6.072xl05
6.289xlOs
6.140xl05
4.402xl05
4,08.2xl05
3.724xl05
L.SOSxlO5
L.192xlOs
U224xl05
3.803xl05
5.840xlOs
3.761xl05
).165xl05
3.259xl05
t
Before
Test Section
% of Input
57.9
67.2
59.7
N.A.
69.5
34.8
60.9
53.4
44.1
53.4
60.5
55.8
43.0
65.2
73.1
68.6
64.0
63.1
60.7
59.4
Test Section
Section I
; of Input
17.7
26.5
32.7
34.2
34.4
28.7
35.4
39.1
35.6
30.2
35.5
39.7
30.0
6.4
12.4
6.7
18.7
24.1
25.6
26.9
Section II
i of Input
8.9
1.6
7.83
12.7
2.13
*' 25.63
1.2
5.23
12.63
24.33
2.0
3.83
12.66
15.03
3.14
28.3
7.63
0.3
33.0
17.0
Section III
% of Input
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
Section IV
?; of Input
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
After Test
Section in air
% of Input
IS.Sct
4.7a
0
N.A.
0
10. 9a
2.5a
2.3a
7,7a
0
2.0a
0.7a
14. 4a
13. 4a
11.4
0 .
9.7a
12. 5a
N.A.
0
•
(see notes on page A-23)
-------
Table A-3. LIQUID MATERIAL BALANCE (continued)
Exp.
No.
70
71
73
74
75
76
77
78
79
80
81
82
83
84
85
86
88
89
90
91
Input
cm 3
7.806X105
7.675xl05
1.386xl06
1.281xl06
3.872xl06
5. .243x10*
4.033xl05
4.438xl05
7.182xl05
6.205xl05
6.207xl05
6.268xl05
1.305x10*
1.278x10*
1.238x10*
1.236x10*
9.418xl05
9.100xl05
1.139x10*
1.147x10*
t
|
*
Before
Test Section
1 of Input
79.2
71.2
30.2
76.0
2.3.
2.9
45.4
57.2
67.4
62.2
53.9
74.9
33.1
14.6
21.3
56.3
20.4
24.2
26.4
38.8
' Test Section
Section I
% of Input
31.5
41.9
46.6
40.9
15.4
25.3
40.4
38.7
37.9
37,3
35,8
25.6
74.2 -
78.8
78.2
53.4
80,2
78.7
80.6
70.8
Section .11
% of Input
1.6
8.5
3.4
0
67,6
60,1
3,3
3.1
0.5
3.4
2.0
0.9
2.4
16.4
9.6
1.0
10.3
8.1
4.3
2.7
Section III
% of Int)ut
0
0.9
2.3
0
6.8
11.9
0
0
0
0
0.1
0
0
0
0
0
0
0.1
0
0
Section IV
% of Innut'
0
0
0
0
0.4
0
0
0
0
0
0
0
0.3
0.1
0
0.1
0.17
0.1
0
After Test
Section in air
% of Incut
N.A.
N.A.
N.A.
N.A.
7.9a
0
10. 9a
l.Oa
0
0
8.2a
0
N.A.
N.A.
N.A.
N.A.
N.A.
N.A.
N.A.
N.A.
(see notes on page A-23)
-------
Table A-3. LIQUID MATERIAL BALANCE (continued)
Bxp.
No,
92
93
94
95
97
98
99
100
Input
cm 3
1.144xl06
1.132xl06
3.892xl06
3.835xl06
3.873xl06
3.880xl06
5.219xlOs
5.448xlOs
Before
Test Section
% of Input
49.7
59.1
2.3
2.1
29.6
44.3
0.9
2.5
Test Section
Section I
of Input
60.9
53.5
29.7
13.3
72.4
57.2
11.6
37.8
section II
' of Input
1,8
0.75
35.5
69.8
0.3
0.17
67.4
51.3
ection III
% of Input
0
0
24.7
15.9
0
0
19.3
4.9
Section IV
% of Input
0
0
12.5
0.6
0
0
1.7
0.4
\fter Test
section in air
% of Input
N.A.
N.A.
N.A.
0
0
0
N.A.
3. la
(see notes on page A-23)
-------
Table A-3. LIQUID MATERIAL BALANCE (continued)
Exp.
No.
101
102
103
104
105
106
107
108
109
110
111
112
113
114
115
Input
cm 3
5.224xl06
5.216xl06
4.526xl05
3.954xl05
4.014xl05
2.473xl05
5.214xl05
L.107xl05
t.635x!05
l.lOSxlO6
L.589xl06
L.607xl06
L.SlOxlO6
L.084xl06
L.088xl06
116 |L.269xl06
117 L.385xlOe
118
119
120
K046xl05
4.007xl05
2.470xl05
Before
Test Section
1 of Input
36.4
50.4
90.3 .
76.8
74.7
61.2
34.3 •
62.4
69.9
85.8
27.1
33.9
21.1
18.4
23.5
17.7
24.4
74.4
75.9
72.8
Test Section
Section I
% of Input
67.5
50.2
44.8
16.3
20.2
14.5
18.1
11.0
23.7
28.5
33.3
58.0
32.4
41.7
34.8
35.4
62.9
14.0
23.8
16.1
Section II
% of Innut
1.6
0.2
0
4.3
0.8
6.7
0
5.5
4.0
0.3
46.9
13.9
52.5
50.9
57.7
60.0
19.9
14.8
0
49.9
Section III
1 of Input
0.09
0
0
0
0
0
0
0
0.2
0
0.4
0
0.2
0.4
0
0.4
0
0
0
0
Section IV
I of Input
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
After Test
Section in air
% of Input
0
0
0
0
0
0
0
0
2.2
0
N.A.
0
N.A.
N.A.
N.A.
N.A.
0
0
0
0
(see notes on page A-23)
-------
Table A-3. LIQUID MATERIAL BALANCE (continued)
to
t-o
5xp.
No.
121
122
123
124
125
126
127
128
130
131
132
133
134
135
136
137
138
140
141
L42
Input
cm3
.088xl05
1.106xl05
.87xl05
3.86xl05
3.86xl05
3.84xl05
3.56xl05
3.56xl05
3.78xl05
3.85xl05
1.252xl06
L.631xl06
L.574xl06
L.529xl06
S.OOOxlO6
7.627xl06
7.629xl06
..842xl06
S.OOSxlO6
5.876xl05
1
«
i
before
Test Section
1 of Input
75.1
63.8
86.5
68.1
91.9
76.7
67.4 •
66.1
63.6
69.4
69.2
17.4
27.9
17.3
16.6
0.8
0.5
90.2
68.2
58.3
Test Section
Section I
\ of Input
20.0
13.7
28.0
28.8
2.6
19.5
26.9
26.8
21.5
8.2
43.6
57.7
74.5
56.0
59.3
22.0
21.8
17.8
39.2
24.7
ection II
of Input
0
2.1
1.75
0
0
0
1.6
0
0
0
0
29.6
1.1
30.4
29.0
61.3
65.4
0
0
0.3
ection III
% of Input
0
0
0
0
0
0
0
0
0
0
0
0.3
0
0.6
0.9
15.7
13.8
0
0
0
ection IV
?: of Input
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0
0.2
0.6
0
0
0
fter Test
ection in air
% of Input
0
0
3.59
0
0
0
4.0
6.8
0
0
N.A.
N.A.
N.A.
N.A.
0.1
0.2
N.A.
N.A.
N.A.
0
i
(see notes on page A-23)
-------
Table A-3. LIQUID MATERIAL BALANCE (continued)
Exp.
No.
143
145
146
147
148
149
150
151
152
153
Input
cm 3
9.025xl05
9.87xl05
9.160xl05
9.415xl05
1.667xl06
3.099xl06
3.006xl06
L.473xl06
J.454xl05
).365xl05
1
j
«i
Before
Test Section
% of Input
56.7
52.9
82.2
71.9
36.8
16.8
30.4'
61.6
87.9
63.3
Test Section
Section I
\ of Input
26.5
23.3
3.0
8.7
29.0
46.9
10.5
43.9
12.4
19.3
Section .11
; of Input
0
0.2
0.2
1.3
22.3
41.4
65.6
3.2
0
2.8
section III
% of Input
0
0
0
0
0
0
0.2
0
0
0
Section IV
% of Input
0
0
0
0
0
0
0
0
0
0
kfter Test
Section in air
% of Input
0
0.1
0
0.1
0.2
N.A.
N.A.
N.A.
N.A.
0
to
C/4
a = Obtained by material balance
3 = Liquid overflowing from previous section
* = If the liquid input is not available (N.A.) liquid flow is given in cm3
-------
Table A-4. PRESSURE DROP DATA
Exp .
No.
i
!
i
1
2
3 '
4
5
6
8
9
11
12
13
14
15
17
18
19
20
21
22
23
24
25
26
27
28
29
30
31
32
33
System
Pressure
mm Hg gauge
N.A.
N.A,
N.A.
N.A.
-24.36.
-26.39
-25.12
-24.61
-23.09
-23.09
-23.09
-24.87
-26.14
-25.12
. -24.80
-23.60 '
-23.50
-26.80
-23.0
-24.4
-24.3 -
-22.1
-21.9
-19.9
-3.7
-20.5
-22.0
-22.0
-22.6
-32.0
Pressure Drop Over Test Section
AP,
dry
cm W . C .
1.7
0.41
3.0
0.41
0.08
0.12
1.5
0.74
0.29
0.12
0.08
0.12
0.04
1.2
1.20
1.1
1,03
0.58
0.29
0.21
N.A.
0.29
0.04
0.04
0.21
0.29
1.3
1.3
1.3
N.A.
A? ^
wet
cm W.C.
1.7
0.41
3.3
0.41
0.12
0.12
1.5
0.82
0.29
0.16
0.08
0.12
0
1.3
1.4
1.5
2.2
1.3
0.62
0.54
0.21
0.45
0.21
0.29
0.54
0.8
1.6
1.8
1.3
N.A.
A-24
-------
Table A-4. PRESSURE DROP DATA (continued)
Exp.
No.
34
35
36
37
38
39
40
41
42
43
45
46
47
48
49
50
51
52
53
54
55
56
58
59
60
61
63
64
65
66
System
Pressure
mm Hg gauge
-22.0
-22.0
-23.0
-23.0
-25.0
-25.5
N.A.
.-23.0
-21.9
-20.5
-23
-23
-23.5
-19
-18
N.A.
-24
-24
-24
-22.8
-23
-23
-20
-20
-22
-21.5
-19
-19
-19
-19
Pressure Drop Over Test Section
AP,
dry
cm w.C.
0.37
0.41
0.12
0.29
1.0
1.1
0.45
0.08
0.04
0.99
0.29
1.0
0.54
0.78
0.61
0.78
0.95
0.66
0.50
1 . 3
0.50
0.79
1.28
0.54
0.83
1.3
1.0
0.74
1.1
0.74
AP «.
wet
cm w.C.
0.37
0.12
0.12
0.54
1.03
1.45
0.62
0.08
0.04
1.4
0.29
1.7
0.87
1.3
0.87
0.87
0.87
0.66
0.50
0.5
0.50
0.79
1.28
0.54
0.91
1.4
1.6
1.1
1.9
1.3
A-25
-------
Table A-4. PRESSURE DROP DATA (continued)
Exp .
No.
67
68
69
70
71
72
73
74
75
76
77
78
79
80
81
82
83
84
85
86
88
89
90
91
92
93
94
95
97
98
99
100
-
System
Pressure
mm Hg gauge
-20
-20
-20
-21.5
-20
-18
-19
-19
-19
-20
-19.8
-22
-23
-22.9
-23
-23
-22
-22
-26
-26.
-19
-20
-20
-20
-20
-20
-20
-20.5
-20
-20
-17 •
-17
Pressure Drop Over Test Section
iPdry
cm W.C.
0.45
1.1
0.78
0.45
8.63
0.45
8.6
0.45
9.1
8.7
2.8
2.4
1.7
2.2
2.7
0.54
0.54
2.7
2.1
0.12
2.7
2.0
1.2
0.540
0.25
0.41
2.7
6.1
0.62
0.21
5.9
3.4
AP «.
wet
cm W.C.
0.78
2.0
1.0
0.54
8.9
0.45
9.29
0.37
4.2
9.5
3.2
2.5
1.8
2.4
2.9
0.54
0.54
3.0
2.2
0.45
2.9
2.0
1.2
0.45
0.25
0.330
2.6
N.A.
0.54
0.12
7.6
5rt !
.0
1
A-26
-------
Table A-4. PRESSURE DROP DATA (continued)
jExp.
No.
101
102
103
104
105
106
107
108
109
110
111
112
113
114
115
116
117
118
119
120
121
122
123
124
125
126
127
128
130
131
System
Pressure
mm Hg gauge
-17.5
-17.5
-19.0
-18.0
-18,0
-17.0
-17.0
-20.0
-20.5
-20.0
-22.0
-22.0
-18.0
-18.0
-17.0
-17.0
-17.0
-23.0
-23.0
-20.0
-20.0
-20.0
-26.0
-25.0
-23.0
-23.0
-20.0
-20.0
-16.0
-17.0
Pressure Drop Over Test Section
AP,
dry
cm w.C.
0.87
0.21
0.21
6.2
2.21
6.2
2.2
6.2
6.1
3.4
5.9
3.5
5.9
6.1
6.2
6.2
3.6
5.9
2.4
6.2
2.4
6.2
1.1
0-.45
0.04
0.29
1.0
0.79
0.29
0.12
APwet
cm w.C.
0.45
0.17
0.21
6.5
2.2
6.5
2.3
6.6
6.5
3.7
6.7
4.2
6.7
6.6
6.7
7.1
4.3
6.4
2.6
7.3
2.5
6.5
1.9
0.87
0.04
0,29
1.9
1.4
0.45
0.12
A-27
-------
Table A-4. PRESSURE DROP DATA (continued)
Exp .
jNo.
132
133
134
135
136
137
138
140
141
142
143
146
147
148
149
150
151
152
153
System
Pressure
mm Hg gauge
-19.0
-17.0
-17.0
-18.6
-23.2
-19.0
-16,8
-17.0
-18.5
-20.0
-20.9
-20.2
-18.5
-14.5
-15.8
-18.0
-16.8
-18.8
-18.7
Pressure Drop Over Test Section
APdry
cm W.C.
0.21
1.]
0.62
1.0
1.0
1.0
1.0
0.12
0.29
0.87
0.62
0.04
0.12
1.2
1.0
0.54
2.60
0.04
5.1
" AP «.
wet
cm W.C.
0.54
1.9
1.5
2.5
2.5
2.9
2.8
0.21
0.62
1.5
0.95
0.04
0.21
1.5
1.2
' 0.79
2.8
0.08
5.4
A-28
-------
APPENDIX B
GLOSSARY
-------
GLOSSARY
COLLECTION EFFICIENCY or OVERALL COLLECTION EFFICIENCY
Collection efficiency or overall collection efficiency
is the mass ratio of net liquid collected in the entrainment
separator to the liquid present in the inlet entrainment.
It can also be expressed as the difference between primary
efficiency and reentrainment. We use either fraction or
percent, as specified in the text or figure.
PENETRATION
Penetration is the mass ratio of drops in the outlet
to that of the inlet of the entrainment separator.
Penetration is also equal to "1 - collection efficiency",
if in fractional form. We use either fraction or percent,
as specified in the text or figure.
PRIMARY COLLECTION
Primary collection is defined as fractional collection
of the drops present in the original entrainment by various
mechanisms which is reported in terms of mass fraction as
an efficiency. Primary efficiency includes only the
collection of drops present in the original entrainment.
The reentrainment of these collected drops or the subsequent
collection of these reentrained drops does not affect the
primary collection efficiency.
REENTRAINMENT
Reentrainment is the mass ratio of drops entering the
gas from the liquid in the entrainment separator to drops
present in the inlet entrainment.
B-l
-------
REENTRAINMENT VELOCITY
Reentrainment velocity is the gas velocity at which
drops are first observed to become reentrained in the gas.
The reentrainment velocity will vary for different kinds of
entrainment separators and different operating conditions.
SECONDARY COLLECTION
Secondary collection refers to drops which are
absorbed after having been reentrained. The mass ratio
of these drops to those present in the inlet entrainment
is the secondary collection efficiency.
B-2
-------
APPENDIX C
REFERENCES
-------
REFERENCES
Alia, P., L. Cravarola, A. Hassid, and E. Pedrouhi.
Liquid Volume Fraction in Adiabatic Two Phase Vertical
Upflow-Round Conduit. CISE Report R-105. 1965.
Anderson, J. D., R. E. Bellinger, and D. E. Lamb. Gas
Phase Controlled Mass Transfer in Two Phase Annular
Horizontal Flow. AIChE Journal. 10;.640, 1964.
Baker, 0. Simultaneous Flow of Oil and Gas. Oil Gas
Journal. 53:185, .1954.
Bell, C. G. and W. Strauss. Effectiveness of Vertical
Mist Eliminators in a Cross Flow Scrubber. Journal
of the APCA. 23:967-9. November 1973.
Bradie, J. K. and A. N. Dickson. Removal of Entrained
Liquid Droplets by Wire Mesh Demisters. Paper 24 in Fluid
Mechanics and Measurements in Two-Phase Flow Systems. (A
joint symposium of the Inst. of Mech. Engr. and the Yorkshire
Branch of the Inst. of Chem. Engr.) 24-25. London. September
1969.
Buerkholz, A..Drop Separation on Wire Fitters. Chemie
Ingenieur Tecknik. 4_2_, 21, 1314-1321, 1970.
Calvert, S. Air Pollution. Stern, A. C. (ed.). 3_, Academic
Press, New York. 1968.
Calvert, S., J. Goldshmid, D. Leith, and D. Mehta.
Scrubber Handbook. Prepared for EPA Contract No. CPA-70-95.
Vol. I and II, 1972.
Calvert,S. and D. Lundgren. Particle Collection in Closed
Packed Arrays. Presented at AIHA. 1970.
Carman, P.C. Fluid Flow Through Granular Beds. Trans.
Inst. Chem. Engr. 15:150, 1937.
Castleman, R.A. Bur. Stand. J. Res. Washington. 6^:369, 1931.
Carpenter, C. L. and D. F. Othmer. Investigation of Wire
Mesh as an Entrainment Separator. AIChE Journal, p. 549, 1955,
Chien, S. F. and W. Ibele, Pressure Drop and Liquid Film
Thickness of Two Phase Annular and Annular-Mist Flows.
ASME Paper. 62-WA170.
Chilton, H. Elimination of Carryover from Packed Towers With
Special Reference to Natural Draught Water Cooling Towers.
Trans. ICE. 30, 235.
C-l
-------
Collier, J. G. and G. F. Hewitt. Data on Vertical Flow
of Air-Water Mixtures in Annular and Dispersed Flow Regions:
Part II, Film Thickness and Entrainment data and Analysis
of Pressure Drop Measurements. Trans. Inst. Chem. Engr,
3£:127, 1961.
Cousins, L. B., W.H. Denton, and G. F, Hewitt. Liquid Mass
Transfer in Annular Two Phase Flow. Exeter. 21-23, June
1965.
Cousins, L. B. and G. F. Hewitt. Liquid Phase Mass Transfer
in Annular Two Phase Flow-Droplet Deposition and Liquid
Entrainment. AERE-R 5657. 1968.
Davis, R. F. Proc. Inst. Mech. Engrs. 149;148, 1940.
Dombrowski, N. and R. P. Frazer. A Photographic Investigation
into the Disintegration of Liquid Sheets. Phil. Trans.
A924, 247, 101, 1954.
Eckert, J. S. Chem. Engr. Prog. 57_, 9, 54, 1961.
Edgerton, H. E., K. J. Germeshansen, and H. E. Grier.
Photography J. 76^:198, 1936.
Page, A. and F. C. Johansen. Proc. Roy Soc. (London).
116A:170, 1927.
Fairs, G. L. Calder-Fox Scrubbers and the Factors
Influencing Their Performance. Trans. I Chem. E. London.
22, 110.
Foust, A. S., L. A. Wenzel, C. W. Clump, L. Maus, and
L. B. Andersen. Principles of Unit Operations. Toppan Company-
Tokyo. 1959.
Garner, F. H., S. R. M. Ellis, and J. A. Laay. The Size
Distribution and Entrainment of Droplets. Trans. Inst.
Chem. Engr. 3_2:222, 1954.
Gill, L. E., G. F. Hewitt, J. W. Hitchon, and P. M. C. Lacy.
Sampling Probe Studies of the Gas Core in Annular Two
Phase Flow: 1, The Effect of Length on Phase and Velocity
Distribution. Chem. Engr. Sci. 1_8:525, 1963.
Golovin, M. N., and A. A. Putnam. Ind. Engr. Chem. Fund.
1_:264, 1962.
Grimison. Trans. Am. Soc. Mech. Engr. 5_9:583-94, 1937.
Grimley, S. S. Trans. Inst. Chem. Engr. (London). 23, 228-235,
1948.
Haenlein, A. Forsch. Ing. Wes. 2^:139, 1931.
C-2
-------
Hall-Taylor, N. S. and R. M. Nedderman. The Coalescence
of Disturbance Waves in Annular Two-Phase Flow. Chem.
Engr. Sci. 23^:551-64, 1963.
Hanratty, T. J. and D. E. Woodmansee. Stability of the
Interface for a Horizontal Air-Liquid Flow. Symposium
on Two Phase Flow. Exeter. Paper Al, 21-23, June 1965.
Hinze, J. 0. Forced Deformation of Viscous Liquid
Globules. Appl. Sci. Res. Ai. 263-72, 1948.
Hinze, J. 0. Critical Speed and Sizes of Globules. Appl.
Sci. Res. Al, 275-88, 1948a.
Houghton, J. G. and W. H. Radford. Trans. Am. Inst. of
Ch. E. 35^:427, 1939.
Hughmark, G. A. Film Thickness Entrainment, and Pressure
Drop in Upward Annular and Dispersed Flow, AIChE Journal.
19, 5, 1062-7, 1973.
Jackson, S. and S. Calvert. AIChE Journal. 1,2:1075, 1966.
Jashnani, I. L. Coalescence and HTU in Foam Fractionation
Columns. Ph.D. Dissertation, U. ov Cincinnati. 1971.
Kitchener, J. A. Foams and Free Liquid Films, in Recent
Progress in Surface Science. Academic Press. New York. 1964.
Kotov, N. A. Influence of Spray Entrainment on Gas Scrubber
Efficiency. Translated by EPA from Vodosnabzheniye i
Sanitaria Technika. 7/.31-4, 1972.
Lane, W. R. Shatter of Drops in Streams of Air. Ind. Engr.
Chem. 43, 1312, 1951.
Lapple, C. E. and C. B. Shepherd. IEC Ind. 32, 605, 1940.
Leith, D. and W. Licht. The Collection Efficiency of
Cyclone Type Particle Collectors - A New Theoretical
Approach. Paper presented at San Francisco meeting of
AIChE. December 1971.
Leva. M. Chem. Engr. Prog. Symp. Ser. No. 10, 50, 51, 1954.
Lobo, W. E., L. Friend, F. Hashmall, and F. A. Zenz.
Trans. AIChE. 41, 693, 1945.
Mercer, T. T. and H. Y. Chow. J. of Coll. and Interface
Sci. 2.7:75-83', 1968.
Miles, J. W. On the Generation of Surface Waves by Shear
Flows. J. Fluid Mech. 3_, 185, 1957.
Moyers, C. G., Jr. Entrainment Separator Selection and
Evaluation. Engr. Res. Report ER-60-1010, Union Carbide.
October 1960.
C-3
-------
Newitt, D. M., N. Dombrowski, and F. H. Knelman. Liquid
Entrainment: 1 The Mechanism of Drop Formation from Gas
or Vapor Bubbles. Trans. Inst. Chem. Engr. 3^, 244, 1954.
Nukiyama, S..and Y. Tarrasawa. Trans. Soc. of Mech. Engr.
(Japan). 4_, 5_, £, 1938-1940.
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Shearer, C. J. and R. M. Nedderman. Pressure Gradient and
Liquid Film Thickness in Cocurrent Upwards Flow of Gas-
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Chem. Engr. Sci. 2£:671-83, 1965.
Shepherd, C. B. and C. E. Lapple. IEC Chem. 31, 1246, 1940.
Sherwood, T. K, and R. L. Pigford. Adsorption and Extraction.
McGraw-Hill. New York. 265, 1952.
Sherwood, T. K., G. H. Shipley, and F. A. L. Holloway. Ind.
Engr. Chem. 3_p_, 765, 1938.
Sorokin, Y. L,, L. .N. Demidova, and N. P. Kuzmin. Principles
of Drop Separation from Vapor or Gas Streams. Translated
from Khimicheskoyi i Nedtinoye Mashinostroyiniye. No. 8,
p. 20, 1968.
Stearman, F. and G. J. Williamson. Spray Elimination in Processes
for Air Pollution Control. Nonhebel, 2nd ed., CRC Press.
Cleveland, 1972.
Steen, D. A. and G. B. Wallis. The Transition from Annular
to Annular-Mist Cocurrent Two-Phase Down Flow, NYO-3114-2, 1964.
Stuhlman, D. Physics, 2_, 457, 1932.
Stumpner, R. Die Werme. 59^, 463, 1936.
C-4
-------
Taheri, M. and S. Calvert. APCA Journal. 18, 240, 1968.
Wallis, G. B. The Onset of Droplet Reentrainment in
Annular Gas-Liquid Flow. General Electric Report. No. 62
GL127, 1962.
Weber, C., Z. Angen. Math. Mech. 2, 139, 1931.
Wicks, M. and A. E. Dukler. Entrainment and Pressure
Drop in Cocurrent Gas Liquid Flow. AIChE Journal. 6, 463,
1960.
Wicks, M. and A. E. Dukler. In Site Measurements of Drop
Size Distribution in Two Phase-Flow: A New Method for
Electrically Conducting Liquids. Paper presented at
International Heat Transfer Conference, Chicago. 1966.
York, 0. H. Performance of Wire Mesh Demisters. Chem.
Engr. Prog. Vol. 50, No. 8, 421, 1954.
York, 0. H. and E. W. Poppele. CEP. 59^ 45, 1963.
Zhivaiking, L. Y. Liquid Film Thickness in Film Type
Units. Int. Chem. Engr. 2, 237, 1962.
C-5
-------
TECHNICAL REPORT DATA
(Please read Instructions on the reverse before completing}
1. REPORT NO.
EPA-650/2-74-119-a
2.
3. RECIPIENT'S ACCESSION NO.
4. TITLE AND SUBTITLE
Entrainment Separators for Scrubbers--Initial
Report
5. REPORT DATE
October 1974
6. PERFORMING ORGANIZATION CODE
7 AUTHOR(sSeymour Calvert, Indrakumar L. Jashnani,
Shuichow Yung, and Samuel Stalberg
8. PERFORMING ORGANIZATION REPORT NO.
9. PERFORMING ORGANIZATION NAME AND ADDRESS
10. PROGRAM ELEMENT NO.
A.P.T. , Inc.
P. O. Box 71
Riverside, CA 92502
1AB013; ROAP 21ACX-086
11. CONTRACT/GRANT NO.
68-02-0637
12. SPONSORING AGENCY NAME AND ADDRESS
EPA, Office of Research and Development
NERC-RTP, Control Systems Laboratory
Research Triangle Park, NC 27711
13. TYPE OF REPORT AND PERIOD COVERED
Initial; Through 6/74
14. SPONSORING AGENCY CODE
15. SUPPLEMENTARY NOTES
16. ABSTRACT
The report gives results of an evaluation of current technology relating to
the separation of entrained drops of liquid from the gas leaving scrubbers. It includes
results of experimental studies of entrainment separator characteristics, as well as
theoretical analysis. Zigzag baffle, knitted mesh, tube bank, packed bed, and cyclone
devices were tested. It reports collection efficiency and reentrainment as related
to drop size. Pressure drop as a function of gas flow rate is also reported.
Mathematical models for primary collection efficiency are satisfactory, but useful
reentrainment models are not yet available. An auxiliary experiment was aimed at
determining reentrainment from liquid sheets under the influence of an air stream.
Work is still in progress on vertical gas flow systems, solids deposition,
mathematical modeling, and the development of improved designs.
7.
KEY WORDS AND DOCUMENT ANALYSIS
DESCRIPTORS
b.lDENTIFIERS/OPEN ENDED TERMS
c. COSATI Field/Group
Air Pollution
Scrubbers
Entrainment
Efficiency
Mathematical Models
Air Pollution Control
Stationary Sources
Entrainment Separators
Collection Efficiency
13B
07A
07D
14A
12A
8. DISTRIBUTION STATEMENT
Unlimited
19. SECURITY CLASS (ThisReport)
Unclassified
21. NO. OF PAGES
318
20. SECURITY CLASS (Thispage)
Unclassified
22. PRICE
EPA Form 2220-1 (9-73)
C-6
------- |