Acetone absorption column

ACETONE ABSORPTION COLUMN
Required: “CONFIDENTIAL” Sh.a. DESIGNED: Arberor MITA Design group“ELMAGERARD”
CALCULATION REPORT
TECHNOLOGICAL AND
CONSTRUCTION PROJECT
EQUIPMENT "ACETONE ABSORPTION
COLUMN"
Nomenclature
1. Inrodiction

Nomenclature
1. Inrodiction
2. Project data
2.1. Problem description
2.2. Packing parameters
2.3. Inlet data
3. Design procedure
3.1. Tower diameter
3.2. Pressure drop
3.3. Diffusion coefficients
3.4. Mass transfer coefficients
3.5. Packing Height
3.6. Operating and equilibrium lines
4.0 Designig of internals
5.0 Determination of nozzles inside diameters
6.0 Mechanical design of absortion column as per ASME S8. Di1/Di2
6.1 Minimum shell thickness
6.2 Torispherical head design
6.3 Shell thicnes at diffrent heights
7.0 Design of support
7.0' Design of support (conventional method)
8.0 Design of flanged joints
9.0 Design an integral radial nozzle centrally located
10.0 Check the design of a radial nozzle in a cylindrical shell
Abstract
In the present work, a packed bed absorption column is designed to remove certain amounts
of acetone contained in a gaseous stream. Four packing types (50-mm metal Hiflow® rings,
50-mm ceramic Pall® rings, 50-mm metal Top Pak® rings rings) are and 25-mm metal VSP®

Abstract
In the present work, a packed bed absorption column is designed to remove certain amounts
of acetone contained in a gaseous stream. Four packing types (50-mm metal Hiflow® rings,
50-mm ceramic Pall® rings, 50-mm metal Top Pak® rings rings) are and 25-mm metal VSP®
considered in order to select the most appropriate one in terms of column dimensions,
pressure drop and mass-transfer results. Several design parameters were determined
including column diameter (D), packing height (Z), overall mass-transfer coefficient (Km) and
gas pressure drop (delta P/Z), as well as the overall number of gas-phase transfer units (NtOG),
overall height of a gas-phase transfer unit (HtOG) and the effective surface area of packing
(ah). The most adequate packing to use for this absorption system constitutes the 25-mm
metal VSP® rings, since it provided the greatest values of Km , ah , as well as the lowest values
of both Z and HtOG , meaning that it will supply the higher mass-transfer conditions with the
lowest column dimensions, but it does not comply with the requirement of pressure drop delta
P/Z() for this reason the next adequate packing to use for this absorption system is 50-mm
ceramic Pall® rings. The influence of both gas mixture (QG) and solvent (mL') feed flowrates on
D, Z, Km, P/Z, NtOG and HtOG was also evaluated for the four packing considered. The design
methodology was solved using computing software MATHCAD PRIME 3.0.
Nomenclature
ah Effective specific surface area
of packing
m-1
A Absorption factor Dimensionless
Ch Hydraulic factor Dimensionless
CL Mass-transfer factor Dimensionless
CP Hydraulic factor Dimensionless
CSflood CS coefficient at flooding
conditions
m/s
CV Mass-transfer factor Dimensionless
dP Effective particle diameter m
D Tower diameter m
DG Gas-phase diffusion
coefficient
m2/s
DL Liquid-phase diffusion
coefficient
m2/s
e/k Lennard-Jones parameter K
fflood Flooding factor %
Fp Packing factor ft-1
Fr Froude number Dimensionless
G Mass velocity kg/m2.s
GMy Gas molar velocity kmol/m2.s
GMx Liquid molar velocity kmol/m2.s
hL Liquid holdup Dimensionless
H Henry’s constant atm
HtOG Overall height of a gas-phase
transfer unit
m
kG Gas-phase convective mass-
transfer coefficient
kmol/m
kL Liquid-phase convective
mass-transfer coefficient
m/s
Km Overall volumetric mass-
kmol/m3
Kv Volumetric mass-transfer
coefficient
kmol/m3

GMy Gas molar velocity kmol/m
GMx Liquid molar velocity kmol/m
hL Liquid holdup Dimensionless
H Henry’s constant atm
HtOG Overall height of a gas-phase
transfer unit
kG Gas-phase convective mass-
kmol/m2.s
kL Liquid-phase convective
mass-transfer coefficient
m/s
Km Overall volumetric mass-
kmol/m3.s
Kv Volumetric mass-transfer
coefficient
kmol/m3.s
KW Wall factor Dimensionless
m Mass flowrate kg/h
M Molecular weight kg/kmol
n Factor Dimensionless
N Molar flowrate kmol/h
NtOG Overall number of gas-phase
transfer units
Dimensionless
ΔPlimit/Z Maximum pressure drop
permitted
Pa/m
ΔP0/Z Dry pressure drop Pa/m
ΔP/Z Overall pressure drop Pa/m
P Pressure atm
Q Volumetric flowrate m3/h
R Ideal gas constant m3.atm/kmol.K
%R Removal percent %
Re Reynolds number Dimensionless
Sc Schmidt number Dimensionless
T Temperature ºC
T* Factor Dimensionless
v Velocity m/s
vflood Velocity at flooding conditions m/s
V Molar volume cm3/mol
X Flow parameter Dimensionless
x Mole fraction in liquid phase Fraction
y Mole fraction in gas phase Fraction
y* Mole fraction in gas phase in
equilibrium with the liquid
Fraction
Z Parking height m
Greek Symbols
Density kg/m3
μ Viscosity Pa.s
σ Collision diameter Å
σAB Average collision diameter Å
ψ0 Dry-packing resistance
coefficient
Dimensionless
Dimensionless
ΩD Diffusion collision integral Dimensionless
1. Inrodiction
Gas-liquid operations are used extensively in chemical and petrochemical industries for

Dimensionless
D Diffusion collision integral Dimensionless
1. Inrodiction
Gas-liquid operations are used extensively in chemical and petrochemical industries for
transferring mass, heat and momentum between the phases. Among the most important gas-
liquid systems employed nowadays is absorption, defined as a mass transfer operation at
which one or more soluble components contained in a gas phase mixture are dissolved into a
liquid solvent whose volatility is low under process conditions. The absorption process could
be classified as physical or chemical. The physical absorption occurs when the target solute is
dissolved into the solvent, while the chemical absorption takes place when the target solute
reacts with the solvent. The removal efficiency of any physical absorption process will depend
on the physical-chemical properties (density, viscosity, diffusivity, etc.) and feed flowrates of
the gaseous and liquid streams; the type of mass-transfer contact surface (packing or plate);
the operating temperature and pressure (commonly, lower temperatures will favor gas
absorption by the liquid solvent); gas-liquid ratio; contact time between phases; and the
solute concentration at the inlet gas stream. Gas-liquid absorption operations are usually
accomplished in equipment named absorbers.
Absorbers are used to a great extent in industrial complexes and plants to separate and purify
gaseous streams, to recover valuable products and chemicals, as well as for contamination
control. The most common absorber types employed in industry are plate columns, packed
towers, Venturi cleaning towers and spray chambers. Packed towers are widely used for gas-
liquid absorption operations and, to a limited extent, for distillations. A typical packed column
consists of a vertical, cylindrical shell containing a support plate for the packing material, mist
eliminators, as well as a liquid distributing device designed to provide effective irrigation to the
packing The liquid is fed at the top of the column and trickles down through the packed bed,
exposing a large surface to contact the gaseous stream, which is supplied at the bottom of the
tower .
The design approach of a packed-bed absorber usually involves the determination of
geometrical parameters such as tower diameter (D) and packing height (Z), as well as some
other mass-transfer and operational variables such as convective mass-transfer coefficients for
gas and liquid streams; dry and overall pressure drops; as well as overall mass-transfer
coefficient. A well designed packed-bed tower will provide the required mass-transfer contact
between gas and liquid phases, with low pressure drop, small capital and operating costs, and
high removal efficiencies.
At the present work, a packed bed absorber is designed to remove certain amounts of acetone
contained in an air stream. Four different packing types (Pall®, Hiflow®, Top Pak® and VSP®)
were evaluated in order to determine which packing configuration provides the lowest column
dimensions (tower diameter and packing height) as well as the highest mass-transfer
coefficient for this application, without exceeding the maximum allowable pressure drop and
also without affecting the requested removal efficiency. The influence of both liquid solvent
and gas mixture feed flowrates on 4 important process parameters (tower diameter, packing
height, gas pressure drop and overall mass-transfer coefficient) was assessed for the four
packing, while the effect of this two flowrates on two design parameters (overall number of
gas-phase transfer units; NtOG and overall height of a gas-phase transfer unit, HtOG) was also
determined.

Figure 1: Typical layout of a packed bed absorber
2. Project data
A gaseous mixture containing air and acetone, with a molar composition of 99 % air and 1 %
of acetone.The ethanol must be removed by means of a countercurrent absorption process

2. Project data
2.1. Problem description
A gaseous mixture containing air and acetone, with a molar composition of 99 % air and 1 %
of acetone.The ethanol must be removed by means of a countercurrent absorption process
using water as the solvent. The gas mixture will enter the tower at a rate of 4000 m3/h, at 27
ºC (300 K) and 1.7 atm, while the solvent (water) will be supplied at a flowrate of 14000 kg/h
and also at 300 K. The required removal of acetone to a final concentration of 100 mg/m3,
while the maximum pressure drop permitted for the gas stream should not exceed 250 Pa/m
of packed height. It’s desired to design a suited packed-bed absorber working at 70% of
flooding and operating under isothermal conditions.
For this application, four packing types will be evaluated (Figure 2):
1. 50-mm metal Hiflow® rings
2. 50-mm ceramic Pall® rings
3. 50-mm metal Top Pak® rings, and
4. 25-mm metal VSP® rings.
Table 1. Performance and mass-transfer characteristics of the different packing considered
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2.2.1. Hydraulic Parameters

2.2.1. Hydraulic Parameters
8¡ 9@ AB
2
Mass-transfer surface area per unit volume
C¡D 9EF 2GH 50 mm Metal Hiflow rings
C¡I 9BFB 2GH 50 mm Cheramic Pall rings
C¡P 9QR 2GH 50 mm Metal Top Pac rings
C¡S 9FTR 2GH 25 mm Metal VSP rings
8U @ Packing porosity or void fraction
CUD TVEQQ 50 mm Metal Hiflow rings
CUI TVQWX 50 mm Cheramic Pall rings
CUP TVEW 50 mm Metal Top Pac rings
CUS TVEQ 25 mm Metal VSP rings
86§ @ Hydraulic factor
C6§D TVWQY 50 mm Metal Hiflow rings
C6§I BVXXR 50 mm Cheramic Pall rings
C6§P TVWWB 50 mm Metal Top Pac rings
C6§S BVXYE 25 mm Metal VSP rings
86 @ Hydraulic factor
C6 D TV`FB 50 mm Metal Hiflow rings
C6 I TVYYF 50 mm Cheramic Pall rings
C6 P TVYT` 50 mm Metal Top Pac rings
C6 S TVQWF 25 mm Metal VSP rings
8a 9@ AB
¨
Packing factor
50 mm Metal Hiflow rings (ft)

Ca D RF 50 mm Metal Hiflow rings (ft)
Ca I B`F 50 mm Cheramic Pall rings (ft)
Ca P `Y 50 mm Metal Top Pac rings (ft)
Ca S BTR 25 mm Metal VSP rings (ft)
2.2.2. Mass trasfer Parameters
864 @ Mass trasfer factor
C64D BVBYW 50 mm Metal Hiflow rings
C64I BVFFQ 50 mm Cheramic Pall rings
C64P BVXFY 50 mm Metal Top Pac rings
C64S BVXQY 25 mm Metal VSP rings
86 @ Mass trasfer factor
C6D TV`TW 50 mm Metal Hiflow rings
C6I TV`BR 50 mm Cheramic Pall rings
C6P TVXWE 50 mm Metal Top Pac rings
C6S TV`TR 25 mm Metal VSP rings
Figure 1. Schematic drawing of the packed bed absorber operationg condition
Since the absorption system operates at low pressure and temperature (1.7 atm and 300 K,
respectively); the solute gas is very diluted in the liquid phase (that is, the liquid phase can be
catalogued as a dilute liquid solution), the system operates under isothermal conditions and
there is no reaction between the dissolved solute and the solvent, it’s assumed that the
system obeys the Henry’s law, the value of the Henry’s constant for an acetone-water system

Since the absorption system operates at low pressure and temperature (1.7 atm and 300 K,
respectively); the solute gas is very diluted in the liquid phase (that is, the liquid phase can be
catalogued as a dilute liquid solution), the system operates under isothermal conditions and
there is no reaction between the dissolved solute and the solvent, it’s assumed that the
system obeys the Henry’s law, the value of the Henry’s constant for an acetone-water system
operating at 27 ºC is H = 3.933 atm. Thus, the distribution coefficient for the gas-liquid
system (acetone-water system) at 27 ºC and 1.7 atm is H/P = 3.933 /1.7 = 2.314.
2.3. Inlet data
The inlet data necessary to carry out the design calculations are showed below:
Cbc `TTT AA2d
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Volumetric flowrate
C7efg TVTB Mole fraction of aceton
C¢hi 9BTGp AA£¦
2d Outlet concentration of acetone in the gass
C2qr B`TTT A£¦
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Mass flowrate of water
C$efg RWVTW AA¦2
2©
Molecular weight of acetone
C$s BW AA¦2
2©
Molecular weight of water
C$etu FWVEY AA¦2
2©
Combined molecular weight of air
Cvw xEEVEE Acetone removal percent
C¨y€‚ xQT Flooding factor
8AAƒ
„
9FRT AA ¡
2
C… 9FRT AA ¡
2
Maximum pressure drop permitted
C†q EEQVT`Q AA£¦
2d Liquid density of solvent
C‡q 9TVTTTWE ¡ % Liquid viscosity of solvent
C†ceˆ‰ BVBYF AA£¦
2d Gas density at atmospheric preasure
C‡efg 9TVTTTXBY ¡ % Vapor viscosity of acetone at 25 ºC
C‡etu 9TVTTTTBE ¡ % Vapor viscosity of air at 25 ºC
Cefg Q`VBYY AA¢2d
2©
Molar volume of acetone
Cetu F`VEFT AA4
2© Molar volume of air
Cefg XVRY` BTGH‘ 2 Collision diameter of acetone
Collision diameter of air

Cetu XVYFB BTGH‘ 2 Collision diameter of air
C(efg XRVFX ’ e/k parameter for acetone
C(etu EQ ’ e/k parameter for air
Cv TVTTTTWFB AAA92d ¡2
92© ’
Ideal gas constant
C§ XVEXX ¡2 Henry constant for ethanol-water system 25 ºC
C“ FQ ”6 System temperature
C BVQ ¡2 System pressure
C• –A§

FVXB` Distribution coefficient
3. Design procedure
The equations and correlations used to design the packed-bed absorber were taken from
different sources ,considering several aspects such as process operating conditions, mass
transfer characteristics, and packing type.

3. Design procedure
The equations and correlations used to design the packed-bed absorber were taken from
different sources ,considering several aspects such as process operating conditions, mass
transfer characteristics, and packing type.
3.1. Tower diameter
The molecular weight of the gas mixture (MG) was determined applying the equation
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The gas mixture density (ρG) at 27 ºC was determined using the Kay’s method
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Viscosity of the gas mixture (μG) was calculated using the following correlation.
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Where μace and μair values are given in Pa.s.
The amount of ethanol absorbed is;
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The amount of solvent liquid exiting the column is:
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The flow parameter (X), the pressure drop parameter under flooding conditions (Yflood) and the
CS coefficient at flooding conditions (CSflood) were determined according to the equations.
Flow parameter X
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Pressure drop parameter under flooding conditions
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S coefficient at flooding conditions

CS coefficient at flooding conditions
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The gas velocity at flooding conditions (vGflood), the gas velocity (vG), and finally the tower
diameter (D), were calculated by using the following correlations:
Gas velocity at flooding conditions
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Gas velocity
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%
C…cS –9…cy€‚S ÿ€‚ TVEYR A2
%
Tower diameter †

Tower diameter
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9` bc
9… …cD
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C‡I –
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AAA
9` bc
9… …cI
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C‡P –
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9` bc
9… …cP
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C‡S –
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9` bc
9… …cS
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3.2. Pressure drop
Most packed-bed absorbers are designed to safely avoid flooding conditions and also to
operate in the preloading region, with a gas-pressure drop limit of 200 – 400 Pa/m of packed
depth . In this approach, both the gas dry pressure drop ( 0/Z) and overall pressure drop

3.2. Pressure drop
Most packed-bed absorbers are designed to safely avoid flooding conditions and also to
operate in the preloading region, with a gas-pressure drop limit of 200 – 400 Pa/m of packed
depth . In this approach, both the gas dry pressure drop (ΔP0/Z) and overall pressure drop
(ΔP/Z) were determined for the absorption process using well-accepted equations. The liquid
holdup influence was also taken into account, that is, when the packed bed is irrigated, the
liquid holdup causes an increment of the pressure drop. Prior to the determination of both
pressure drops, it was necessary to determine several parameters first. Among those
parameters are included the effective particle diameter (dP) ; the wall factor (KW); the gas-
phase Reynolds number (ReG); the dry-packing resistance coefficient (ψ0); liquid mass velocity
(GL); the liquid velocity (vL); the liquid-phase Reynolds number (ReL); liquid-phase Froude
number (FrL) ; the ratio ah/a; the effective specific surface area of packing (ah); and, finally,
the liquid holdup (hL).
Effective particle diameter
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Wall factor
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Gas-phase Reynolds number
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Dry-packing resistance coefficient
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Liquid mass velocity
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9… ‡D
u `VWRR AA£¦
92u %
C‰qI –AAA
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9… ‡I
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9… ‡P
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92u %
C‰qS –AAA
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9… ‡S
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Liquid velocity
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%
C…qI –AA
‰qI
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%
C…qP –AA
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%
C…qS –AA
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%
Liquid-phase Reynolds number

Liquid-phase Reynolds number
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Liquid-phase Froude number
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Ratio ah/a
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Effective specific surface area of packing
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2
H
ŒLiquid holdup

Liquid holdup
C)qD –99BF
˜
h
™
AA
a'qD
v(qD
d
i
e
ŒH
d
ŠD
Œu
d TVBQX
C)qI –99BF
˜
h
™
AA
a'qI
v(qI
d
i
e
ŒH
d
ŠI
Œu
d TVBE`
C)qP –99BF
˜
h
™
AA
a'qP
v(qP
d
i
e
ŒH
d
ŠP
Œu
d TVBY
C)qS –99BF
˜
h
™
AA
a'qS
v(qS
d
i
e
ŒH
d
ŠS
Œu
d TVFEX
The gas dry pressure drop per meter of packing height (ΔP0/Z) was determined according to
the following correlation:
C…‚uhD –9iD AAAA
99¡D ŽcD
u †g
9F ’sD UD
d QFVTXB AA ¡
2
C…‚uhI –9iI AAAA
99¡I ŽcI
u †g
9F ’sI UI
d BBBVY` AA ¡
2
C…‚uhP –9iP AAAA
99¡P ŽcP
u †g
9F ’sP UP
d EXVXT` AA ¡
2
C…‚uhS –9iS AAAA
99¡S ŽcS
u †g
9F ’sS US
d BYXVFYF AA ¡
2
The gas overall pressure drop per meter of packing height (ΔP/Z) can be finally calculated:
C…gˆD –9…‚uhD
˜
h
h™
˜
h
™
AAA
UD
fUD )qD
d
i
e
Hoj
(
ŒŒ‘’“
u‘‘
d
i
ie
BFEVQTF AA ¡
2
C…gˆI –9…‚uhI
˜
h
h™
˜
h
™
AAA
UI
fUI )qI
d
i
e
Hoj
(
ŒŒ‘’”
u‘‘
d
i
ie
BEYVFYF AA ¡
2
C…gˆP –9…‚uhP
˜
h
h™
˜
h
™
AAA
UP
fUP )qP
d
i
e
Hoj
(
ŒŒ‘’•
u‘‘
d
i
ie
BQEVYTX AA ¡
2
C…gˆS –9…‚uhS
˜
h
h™
˜
h
™
AAA
US
fUS )qS
d
i
e
Hoj
(
ŒŒ‘’–
u‘‘
d
i
ie
XTQVRXB AA ¡
2
Gas-phase diffusion coefficient: The theory describing diffusion processes in binary gas
mixtures at low to moderate pressures has been studied extensively in recent years, and is
well developed nowadays. Since the absorption process is a binary gas system taking place at

Gas-phase diffusion coefficient: The theory describing diffusion processes in binary gas
mixtures at low to moderate pressures has been studied extensively in recent years, and is
well developed nowadays. Since the absorption process is a binary gas system taking place at
low-pressure, the gas-phase diffusion coefficient can be estimated using the Wilke and Lee
correlation :
Gas diffusion coefficient
8—c AAAAAAAAA
9
˜
h
h™
fXVTX AAATVEW
˜˜˜˜u $™š
d
i
ie
BTGd “
Œd
u
999 ˜˜˜˜u $™š ™š
u ›œ
Molecular weight of the gas mixture
C$™š –F
˜
h
™
—AAB
$efg
AAB
$etu
d
i
e
GH
TVTXE AA£¦
2©
Colision diameter of the micture
C™š –AAAA
—efg etu
F
˜™ 9XVREX BTGH‘de 2
Difusion colision diameter integral of the mixture
8›œ ———AAABVTYTXY
“w‘oHjH‘ AAATVBEXTT
( y‘opždj Ÿ AAABVTXRWQ
( yHoju€€ Ÿ AAABVQY`Q`
( ydoz€pHH Ÿ
C“w –AAAA“
˜˜˜˜˜˜˜u 9(efg (etu
RVBX`
C›œ –———AAABVTYTXY
“w‘oHjH‘ AAATVBEXTT
( y‘opždj Ÿ AAABVTXRWQ
( yHoju€€ Ÿ AAABVQY`Q`
( ydoz€pHH Ÿ TVWXE
C—c 99BTVW BTG AA2u
%
Liquid-phase diffusion coefficient: Compared with the kinetic theory behind the gases
behavior, which is well developed and available today, the theoretical basis of the internal
structure of liquids and their transport characteristics are still insufficient to permit a rigorous
treatment. Usually, liquid diffusion coefficients are several orders of magnitude smaller than
gas diffusivities, and depend mostly on concentration profiles due to changes in viscosity, as
well as some changes in the degree of ideality of the solution. To determine the liquid-phase
diffusion coefficient in binary systems for solutes transport to aqueous solutions, the Hayduk
and Minhas correlation was used:
Liquid-phase diffusion coefficient
8—q AAAAAAAAAAAA
99BVFR BTGz ˜™ fgˆ‹
G‘oH€ TVFEFde “Hoju ‡q
¡
BTTT
8¥ AAEVRW
gˆ‹
BVBF C—q 99BVBW BTG€ AA2u
%
To determine the mass transfer coefficients for both phases, two correlations were used which
were obtained from an extensive study, that involved measurement and correlation of mass-
transfer coefficients for 31 different binary and ternary systems, equipped with 67 different
types and sizes of packings, in columns of diameter ranging from 6 cm to 1.4 m.

To determine the mass transfer coefficients for both phases, two correlations were used which
were obtained from an extensive study, that involved measurement and correlation of mass-
transfer coefficients for 31 different binary and ternary systems, equipped with 67 different
types and sizes of packings, in columns of diameter ranging from 6 cm to 1.4 m.
Gas-phase convective mass-transfer coefficient (kG):
8£c 9999TVBXT` 6D
˜
h
™
AAA
9—c
9v “
d
i
e
˜
h
h™
AAAAA
¡D
˜˜˜˜˜˜˜˜˜˜u 9UD ˜™ fUD )qde
d
i
ie
˜
h
™
AA
v(c
’s
d
i
e
Œd
p
#¢c
Œu
d
ScG - Shmit number for gas phase
C#¢c –AAA
‡g
9†g —c
TVWWW
C£cD –99999TVBXT` 6D
˜
h
™
AAA
9—c
9v “
d
i
e
˜
h
h™
AAAAAA
¡D
˜˜˜˜˜˜˜˜˜˜˜u 9UD ˜™ fUD )qDde
d
i
ie
˜
h
™
AA
v(cD
’sD
d
i
e
Œd
p
#¢c XV`YQ AA2©
92u %
C£cI –99999TVBXT` 6I
˜
h
™
AAA
9—c
9v “
d
i
e
˜
h
h™
AAAAAA
¡I
˜˜˜˜˜˜˜˜˜˜˜u 9UI ˜™ fUI )qIde
d
i
ie
˜
h
™
AA
v(cI
’sI
d
i
e
Œd
p
#¢c XVXWX AA2©
92u %
C£cP –99999TVBXT` 6P
˜
h
™
AAA
9—c
9v “
d
i
e
˜
h
h™
AAAAAA
¡P
˜˜˜˜˜˜˜˜˜˜˜u 9UP ˜™ fUP )qPde
d
i
ie
˜
h
™
AA
v(cP
’sP
d
i
e
Œd
p
#¢c XVFRX AA2©
92u %
C£cS –99999TVBXT` 6S
˜
h
™
AAA
9—c
9v “
d
i
e
˜
h
h™
AAAAAA
¡S
˜˜˜˜˜˜˜˜˜˜˜u 9US ˜™ fUS )qSde
d
i
ie
˜
h
™
AA
v(cS
’sS
d
i
e
Œd
p
#¢c XVRX` AA2©
92u %
Liquid-phase convective mass-transfer coefficient (kL):
C£qD –99TVQRQ 64D
˜
h
™
AAAA
99—q ¡D ŽqD
9UD )qD
d
i
e
‘oj
˜™ 9`VER BTGjde A2
%
C£qI –99TVQRQ 64I
˜
h
™
AAAA
99—q ¡I ŽqI
9UI )qI
d
i
e
‘oj
˜™ 9`VWWR BTGjde A2
%
C£qP –99TVQRQ 64P
˜
h
™
AAAA
99—q ¡P ŽqP
9UP )qP
d
i
e
‘oj
˜™ 9RV`BW BTGjde A2
%
C£qS –99TVQRQ 64S
˜
h
™
AAAA
99—q ¡S ŽqS
9US )qS
d
i
e
‘oj
˜™ 9RVYFR BTGjde A2
%
where:
CL – Mass transfer factor
a – Mass transfer surface area per unit volume
3.5. Packing Height
In those systems handling dilute solutions and when Henry’s law applies, is very usual and
convenient to work with overall mass-transfer coefficients in order to calculate the packing
height (Z), which can be determined by the following expression:

where:
– Mass transfer factor
a – Mass transfer surface area per unit volume
3.5. Packing Height
In those systems handling dilute solutions and when Henry’s law applies, is very usual and
convenient to work with overall mass-transfer coefficients in order to calculate the packing
height (Z), which can be determined by the following expression:
8„ —§ˆ¢c £ˆ¢c
where:
HtOG – Overall height of a gas-phase transfer units
NtOG – Overall number of gas-phase transfer units
Prior to determine the values of HTU and NTU, it will be necessary to calculate several
parameters first, which are the inlet gas molar velocity [GMy(1)]; the outlet gas molar velocity
[GMy(2)]; the average molar gas velocity (GMy); the inlet liquid molar velocity [GMx(2)]; the outlet
liquid molar velocity [GMx(1)]; the absorption factor at the bottom [A(1)] and top [A(2)] of the
column ; the geometric average of the absorption factor (A) ; the ethanol molar composition
of outlet gas [yeth(2)] ; the volumetric gas-phase (KvG) and liquid-phase (KvL) mass-transfer
coefficients ; the overall volumetric mass-transfer coefficient (Km); the overall height of a gas-
phase transfer unit (HtOG) ; the overall number of gas-phase transfer units (NtOG); and finally
the packing height (Z).
Inlet gas molar velocity
8¤¥h¦ AAA
` §c
9… ü
Gas molar flow rate
C§c –AAA
9bc †g
$c
˜™ 9FVQRE BTjde AA2©
)'
Inlet gas molar velocity
C¤¥h¦D –AAA
` §c
9… ¨D
u E`VYTW AA2©
92u %
C¤¥h¦I –AAA
` §c
9… Ï
u RQVFRB AA2©
92u %
C¤¥h¦P –AAA
` §c
9… ¨P
u BTTVRWE AA2©
92u %
C¤¥h¦S –AAA
` §c
9… ¨S
u YYVRQW AA2©
92u %
8¤¥h© AAAAA
` ˜™ f§c §gˆ‹elmde
9… ü
Molar flow of ethanol absorbed
C§efgelm –99AAA
9bc †g
$c
7efg vw TVQYY AA2©
%
Outlet gas molar velocity

Outlet gas molar velocity
C¤¥h©D –AAAAA
` ˜™ f§c §efgelmde
9… ¨D
u EXVYYF AA2©
92u %
C¤¥h©I –AAAAA
9… Ï
u RYVYQE AA2©
92u %
C¤¥h©P –AAAAA
9… ¨P
u EEVRWX AA2©
92u %
C¤¥h©S –AAAAA
9… ¨S
u YRVEBX AA2©
92u %
Average molar gas velocity
C¤¥hD –AAAAA
—¤¥h¦D ¤¥h©D
F
E`VBXR AA2©
92u %
C¤¥hI –AAAAA
—¤¥h¦I ¤¥h©I
F
RYVEYR AA2©
92u %
C¤¥hP –AAAAA
—¤¥h¦P ¤¥h©P
F
BTTVTWY AA2©
92u %
C¤¥hS –AAAAA
—¤¥h¦S ¤¥h©S
F
YYVF`Y AA2©
92u %
Inlet liquid molar velocity
8¤¥ª© AAA
` §q
9… ü
Inlet liquid molar flow rate
C§q –AA
2q
$s
FBWVRFF AA2©
%
C¤¥ª©D –AAA
` §q
9… ¨D
u FYEVQB AA2©
92u %
C¤¥ª©I –AAA
` §q
9… Ï
u BYXVFBX AA2©
92u %
C¤¥ª©P –AAA
` §q
9… ¨P
u FWYVQYB AA2©
92u %
C¤¥ª©S –AAA
` §q
9… ¨S
u BWEVWT` AA2©
92u %

Outlet liquid molar velocity C¤¥ª¦D –AAAAA
` ˜™ f§q §efgelmde
9… ¨D
u FYWVQY` AA2©
92u %
C¤¥ª¦I –AAAAA
9… ¨I
u BYFVY`B AA2©
92u %
C¤¥ª¦P –AAAAA
9… ¨P
u FWRVQRR AA2©
92u %
C¤¥ª¦S –AAAAA
9… ¨S
u BWEVBXW AA2©
92u %Absorption factor at the bottom
–• FVXB` C«¬D –AAAA
®¯¬D
9®h¬D •
BVFFW
C«¬I –AAA
®¯¬I
9®h¬I •
BVFFW
C«¬P –AAA
®¯¬P
9®h¬P •
BVFFW
C«¬S –AAA
®¯¬S
9®h¬S •
BVFFW
Absorption factor at the top
C«°D –AAAA
®¯°D
9®h°D •
BVF`R
C«°I –AAA
®¯°I
9®h°I •
BVF`R
C«°P –AAA
®¯°P
9®h°P •
BVF`R
C«°S –AAA
®¯°S
9®h°S •
BVF`R
Geometric average of the absorption factor
C«D –AAAA
—«¬D «°D
F
BVFXY
C«I –AAAA
—«¬I «°I
F
BVFXY
C«P –AAAA
—«¬P «°P
F
BVFXY
C«S –AAAA
—«¬S «°S
F
BVFXY
Ethanol molar composition of outlet gas

Ethanol molar composition of outlet gas
C7efgr –97efg ss fB vwtt 9BT BTG±
Volumetric gas-phase mass-transfer coefficient
C’ScD –9£cD ¡‹D FW`VFBR AA2©
92d %
C’ScI –9£cI ¡‹I `FRVRQF AA2©
92d %
C’ScP –9£cP ¡‹P FXBVQRY AA2©
92d %
C’ScS –9£cS ¡‹S QYXVW`R AA2©
92d %
Volumetric liquid-phase mass-transfer coefficient
C¢ –AA
†q
$s
˜™ 9RVRXE BTpde AA2©
2d
C’SqD –99£qD ¡‹D ¢ FF`VQ`` AA2©
92d %
C’SqI –99£qI ¡‹I ¢ X`TVX`F AA2©
92d %
C’SqP –99£qP ¡‹P ¢ FBXVQEF AA2©
92d %
C’SqS –99£qS ¡‹S ¢ YQXVXTW AA2©
92d %
Overall volumetric mass-transfer coefficient
C’‰D –AAAAAB
—AAB
’ScD
AA•
’SqD
QFVXEW AA2©
92d %
C’‰I –AAAAAB
—AAB
’ScI
AA•
’SqI
BTEVXF AA2©
92d %
C’‰P –AAAAAB
—AAB
’ScP
AA•
’SqP
YYVTYY AA2©
92d %
C’‰S –AAAAAB
—AAB
’ScS
AA•
’SqS
FBTVQXW AA2©
92d %
Overall height of a gas-phase transfer unit

Overall height of a gas-phase transfer unit
C§ˆ²cD –AA
®hD
’‰D
BVX 2
C§ˆ²cI –AA
®hI
’‰I
TVRFB 2
C§ˆ²cP –AA
®hP
’‰P
BVRBR 2
C§ˆ²cS –AA
®hS
’‰S
TVXB` 2
Product streams and composition
C³ –´c ˜™ 9FVQRE BTjde AA2©
)'
C7³ TVTB AA2©
2©
Cni T
Molar ratio of acetone in the input gas
Cp³ –AA
7³
fB 7³
TVTB AA2©
2©Molar flowrate of carier gas
C –9³ ˜™ fB p³de ˜™ 9FVQXF BTjde AA2©
)'Concentration of acetone in the output gas
C7i –9¢hi AAA
$etu
9$efg †g
˜™ 9FV`QB BTGjde AA2©
2©
Molar ratio of acetone in the output gas
Cpi –AA
7i
fB 7i
˜™ 9FV`QB BTGjde AA2©
2©
Molar ratio of acetone in the output liquid (equilibrium case)
Cn³‰ –A
p³
•
TVTT` AA2©
2©Molar flowrate of liquid
C4 –AA
2qr
$s
˜™ 9QVQQW BTjde AA2©
)'
Molar ratio of acetone in the output liquid
Cn³ –—ni A
4
˜™ fp³ pide TVTT` AA2©
2©
Diferece in molar ratios of acetone in the output gas
Cƒp³ –fp³ 9• n³ TVTTF AA2©
2©Diferece in molar ratios of acetone in the input gas
Cƒpi –—pi 9• ni ˜™ 9FV`QB BTGjde AA2©
2©
Number of gas-phase transfer units

Number of gas-phase transfer units
C´ˆ²cD –9AAAA
fp³ pi
fƒp³ ƒpi
qr
˜
h
™
AA
ƒp³
ƒpi
d
i
e
FXVFTB
C´ˆ²cI –9AAAA
fp³ pi
fƒp³ ƒpi
qr
˜
h
™
AA
ƒp³
ƒpi
d
i
e
FXVFTB
C´ˆ²cP –9AAAA
fp³ pi
fƒp³ ƒpi
qr
˜
h
™
AA
ƒp³
ƒpi
d
i
e
FXVFTB
C´ˆ²cS –9AAAA
fp³ pi
fƒp³ ƒpi
qr
˜
h
™
AA
ƒp³
ƒpi
d
i
e
FXVFTB
Packing height
C„ –9´ˆ²cD §ˆ²cD XTVBYQ 2
C„ –9´ˆ²cI §ˆ²cI BFVTE 2
C„ –9´ˆ²cP §ˆ²cP XRVB`W 2
C„ –9´ˆ²cS §ˆ²cS QVFEX 2
3.6. Operating and equilibrium lines
The operating line will be elaborated using the following data:
Mole fraction of acetone in inlet gas mixture [yace(1)] = 0.01
Mole fraction of acetone in outlet gas mixture [yace(2)] = 9FV`QB BTGj
Mole fraction of acetone in inlet liquid [xace(2)] = 0.
Mole fraction of acetone in outlet liquid [xace(1)] = TVTTX
The operating line points
C7ww TVTB
9FV`QB BTGj
µ
¶·
¸
¹º
C»ww TVTTX
T
µ
¶·
¸
¹º
The equilibrium line for the absortion system
8887w 9• » 9A§

» 9FVXB` »
C7wss»tt 9FVXB` »
The operating and equilibrium line for the absortion system

The operating and equilibrium line for the absortion system
¼½¼¼¾
¼½¼¼¿
¼½¼¼À
¼½¼¼Á
¼½¼¼Â
¼½¼¼Ã
¼½¼¼Ä
¼½¼¼Å
¼
¼½¼¼Æ
¼½¼Æ
ÄÇÆ¼ÈÉ ¼½¼¼Æ ¼½¼¼¾ ¼½¼¼¾ ¼½¼¼¾ ¼½¼¼¿ ¼½¼¼¿ ¼½¼¼À¼ ÀÇÆ¼ÈÉ ¼½¼¼À
»
»ww
9FVXB` »
7ww

(Designing of internals is proprietary knowledge for me so I will not give you a step by step guide how
to design or select internals, instead I will give you a the theoretical fundamentals on how to do i )
Theoretical fundamentalsThe basis of any distributor design is the exact knowledge of the
discharge behaviour of liquids from ground holes and lateral rectangular slots or triangular

(Designing of internals is proprietary knowledge for me so I will not give you a step by step guide how
to design or select internals, instead I will give you a the theoretical fundamentals on how to do it)
Theoretical fundamentalsThe basis of any distributor design is the exact knowledge of the
discharge behaviour of liquids from ground holes and lateral rectangular slots or triangular
notches. The following refers to circular ground orifices but can be applied analogously to
rectangular slots or triangular notches as well. The fundamentals of the discharge behaviour
of fluids out of circular openings stretch back to the year 1644, where Torricelli developed
Equation 1.Michael Schultes, Werner Grosshans, Steffen Müller and Michael Rink, Raschig
GmbH, Germany, present a modern liquid distributor and redistributor design.
ÊËÌÍ ÎÎÎÎÎÏ
ÐÑ Ò Ó
Equation 1 describes the theoretical discharge velocity of liquids, wth, from orifices as a
function of the gravitational acceleration, g, and the liquid head above the orifice, h. If one
multiplies this theoretical velocity, wth, by the cross-sectional area of a hole, Ah, and the
number of discharge holes of a liquid distributor, n, then one achieves the theoretical total
volume rate, , which can flow out of a liquid distributor (Equation 2).
ÊÔÌÍ ÐÐÕÍ Ö ÎÎÎÎÎÏ
ÐÑ Ò Ó
Equation 2 applies under ideal conditions, i.e. assuming that the flow through the hole imparts
no resistance to the flow of liquid. But in reality, streamlines of different velocities are formed
due to the sharp edged holes which cause deflection of the liquid jet flow (Figure 1). For
describing the flow behaviour of the liquid jet flow, one has to interpret two effects. First the
jet contraction and second the jet velocity. The orientation of the streamlines causes the jet of
liquid to contract when it leaves the ground hole. This effect can be described mathematically
by a contraction coefficient, CC. Friction losses, caused by shearing forces of the fluid,
influence the velocity of the jet of liquid when it issues through the hole and can be described
mathematically by a velocity coefficient, CV. The coefficients depend on the liquid head, the
hole geometry and the physical properties of the liquid.The product of the contraction
coefficient, CC, and the velocity coefficient, CV, results in the discharge coefficient, CD = CC ·
CV, which describes the difference between the effective volume rate, , and the theoretical
value,. Only the discharge coefficient, CD, can be derived from experimental investigations
directly.
ÊÔ ÐÐÐ×Ø ÕÍ Ö ÎÎÎÎÎÏ
ÐÑ Ò Ó
The discharge coefficient, CD, is described in the literature according to Table1 as a constant
value in function of the hole geometry only.
Actual discharge behaviour
If one describes the flow through a hole on the basis of an energy balance, equilibrium can be
set up according to Equation 4. The inflowing volume, , acts on the hole with the potential
energy (?L -?V)gh while the out flowing volume rate leaves the hole as a jet flow with the
kinetic energy (?Lw2/2). The energy of the jet leaving the hole is less than that of the

Actual discharge behaviour
If one describes the flow through a hole on the basis of an energy balance, equilibrium can be
set up according to Equation 4. The inflowing volume, , acts on the hole with the potential
energy (?L -?V)gh while the out flowing volume rate leaves the hole as a jet flow with the
kinetic energy (?Lw2/2). The energy of the jet leaving the hole is less than that of the
inflowing liquid since the contraction and friction loss of the jet has consumed energy
characterised by in Equation 4.
ÊÐÐÐÙÚ ÛÜÝ ÜÞßà Ò Ó Ô áÐÐâ
ÜÝ
Ñ
ãÏ Ô ä
In Equation 4, ρL describes the liquid density and ρV the gas density; g describes the
gravitational acceleration, h describes the liquid head above the hole, and w describes the
current velocity of the jet. By including Equation 3 in Equation 4, Equation 5 follows for the
coefficient of discharge CD. The second term of the right side of Equation 5 describes the
energy consumption E divided by the potential energy for a certain liquid head above the hole
h and for a volume flow rate V. The energy consumption E tends to zero for low flow rates.
Consequently, the coefficient of discharge CD tends to unity for low flow rates in case the gas
density can be neglected compared to the liquid density.
Ê×Ø
ÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÏ
Ûâââ
ÛÜÝ ÜÞ
ÜÝ
ââââä
ÐÐÐÜÝ Ò Ó Ô
Systematic investigations have shown that the discharge coefficient, CD = CC CV is a variable
which is dependent on several influencing parameters. An expression for the energy
consumption E was evaluated that allows an accurate prediction of the coefficient of discharge
CD.Following the influencing parameters on the overall coefficient of discharge, CD, will be
discussed in detail (Figure 2). Figure 2 shows measurement points for the discharge
coefficient, CD, for water at ambient temperature as a function of the liquid head, h, for
various hole diameters, d. It also shows the constant CD = 0.62 recorded in the literature for
holes whose dimensions are larger than the depth of the hole.12 It can be clearly seen that
the discharge coefficients only approximate the value given in the literature if the liquid head
is great and hole diameters large. With decreasing liquid head, the discharge coefficient rises
significantly with the result that the discharge behaviour with small liquid head deviate more
favourable from discharge behaviour according to Table 1 than with large liquid head. This can
be explained by the fact that as the liquid head decreases, the horizontal velocity component
decreases and therefore a reduction of the contraction of the jet occurs. Figure 2 also shows
that with decreasing hole diameter, the discharge coefficient rises, i.e. the contraction is
reduced by the counteraction of the horizontal velocity components. In case of small liquid
heads, tensile forces of the jet are also transferred even into the hole cross-section, with the
result that the liquid is drawn out of the opening and, if the liquid level is calm, a vortex is
formed. This effect is more marked in the case of large hole geometries than in the case of
small hole geometries, as can be seen from the steeper curves in Figure 2 at low liquid heads.
The relationships described only apply if the influence of the surface tension is negligible. For
instance, in case of small holes and liquids with a high surface tension, a drop of liquid is
formed beneath the hole, preventing the fluid from flowing out. Further factors that are
influencing the coefficient of discharge but not discussed here are physical properties of the
fluid (density, viscosity), elevation and orientation of holes, overflow velocity and ratio of hole
diameter to deck thickness

instance, in case of small holes and liquids with a high surface tension, a drop of liquid is
formed beneath the hole, preventing the fluid from flowing out. Further factors that are
influencing the coefficient of discharge but not discussed here are physical properties of the
fluid (density, viscosity), elevation and orientation of holes, overflow velocity and ratio of hole
diameter to deck thickness
The dimensioning of liquid distributors
In the dimensioning of liquid distributors, not only the discharge coefficient but also other
design determining criteria have to be taken into account. In that manner, first the minimum
liquid head, hmin , above a discharge hole of a liquid distributor must be determined. Here the
flow velocity of the liquid in the distributor troughs is of decisive importance since flow
gradients occur due to wall friction and lead to significant differences in height, particularly in
case of low liquid heads. If these minimum liquid heads are not attained, considerable
maldistribution of a distributor can occur, as described as follows. The minimal liquid level in a
distributor Feeding liquid into a mass transfer column generally takes place via a feed pipe
which first leads the liquid into a preliminary distribution trough called a parting box.
Afterwards, the liquid reaches the individual distributor troughs lying below and flows in the
troughs towards the column wall. Drag, due to wall friction, causes liquid head gradients to
form in the distributor troughs, which has to be taken into account, particularly with low liquid
levels. This is illustrated in Figure 3.
For liquid flow in open trough systems the difference in height, ∆h, taking place along the
trough can be calculated using Equation 6. It is a function of the wall friction factor, λ,
hydraulic diameter, dh, gravitational acceleration, g, and flow velocity. Figure 3 shows the
result of such a calculation for a trough with a width of b = 100 mm for various flow velocities
as a function of the liquid head, h.
ÊåÓ Ðæ âç
èÍ
ââãÏ
Ñ Ò
ÊÊèÍ é âÕ
ê
âââÐÓ ë
áë Ñ Ó
The following recommendations for the minimum liquid level, hmin , in distributor troughs
can be derived from the observations of Figure 3. The minimum liquid head should not fall
short of a value of 25 - 35 mm and at the same time the flow velocity in the troughs should
be limited to a maximum of 0.5 m/s. Furthermore, the liquid head should be equivalent to at
least twice the hole diameter in order to avoid vortex formation above the hole (Equation 7).
The overall height of a distributor
The overall height of a liquid distributor is first defined by the required liquid loading range.
To this height, the hydrostatic pressure occurring as the result of the pressure drop of the
gas phase as this passes the troughs of the distributor must be added. Furthermore,

The following recommendations for the minimum liquid level, hmin , in distributor troughs
can be derived from the observations of Figure 3. The minimum liquid head should not fall
short of a value of 25 - 35 mm and at the same time the flow velocity in the troughs should
be limited to a maximum of 0.5 m/s. Furthermore, the liquid head should be equivalent to at
least twice the hole diameter in order to avoid vortex formation above the hole (Equation 7).
The overall height of a distributor
The overall height of a liquid distributor is first defined by the required liquid loading range.
To this height, the hydrostatic pressure occurring as the result of the pressure drop of the
gas phase as this passes the troughs of the distributor must be added. Furthermore,
additional height is necessary if a foaming system is present and if a noticeable gas rate is
injected into the liquid at the liquid feed point. The latter applies particularly in the case of
high pressure systems if the degassing of the liquid is markedly restricted due to the small
differences in density between the gas and the liquid. Furthermore, wave formation has to
be taken into account in the case of flowing liquids.
Liquid loading range
By converting Equation 3, Equation 8 can be obtained which defines the necessary extra
height, ?h1, of a liquid distributor resulting from a required loading range.
ÊåÓì Ð
Ù
í
íÚ
ÛÐ
Ù
í
Ú
ââ
Ôîïð
Ôîñò
ß
ó
à
Ï Ù
í
Ú
âââ
Ð×ô Óîñò
Ð×ô Óîïð
ß
ó
à
Ï
ç
ß
ó
óà
Óîñò
Reprinted from HydrocarbonEnginEEringJanuary2009 www.hydrocarbonengineering.com
When calculating the necessary extra height, Figure 2 must be taken into account showing
that the discharge coefficient, CD, provides larger values with the lower liquid head than with
the higher liquid head. This yields greater overall heights than if a constant discharge
coefficient is assumed.
Gas phase pressure drop
The pressure drop, which the gas flow undergoes when it passes through the narrowed
distributor cross-section, can be calculated by Equation 9. ? is the drag coefficient for the
sudden narrowing and expansion of flows, Fv is the gas capacity factor in the column, AC is
the cross-sectional area of the column and AD is the free cross-sectional area of the
distributor.
Êåõ Ðâö
Ñ
ââ
Õ÷
Ï
ÕôÏ øù
Ï Êåõ ÐÐÙÚ ÛÜÝ ÜÞßà Ò åÓú
Ï
This pressure drop causes a rise in hydrostatic pressure to the head of liquid in the distributor,
which can be described with the aid of Equation 10. By equating Equation 9 and Equation 10,
Equation 11 can be obtained for the description of the second portion, ?h2, for the overall
height of a distributor.
ÊåÓú Ðââââç
ÐÙÚ ÛÜÝ ÜÞßà Ò
âö
Ñ
ââ
Õ÷
Ï
ÕôÏ øù
Ï
Foaming systemIn the case of a foaming system, the foam will be built up in particular in
those areas in which a marked gas injection into the liquid takes place. This applies
particularly to the transfer of the liquid from the feed pipe into the parting box, since
relatively large quantities of liquid are transferred per transition point. Since the description of
the foaming behaviour is very complex, it is advisable to use the foam or system factor, ?,
which is described in the literature.13 This empirical factor is known for numerous mass
transfer tasks and has to be taken into account by Equation 12, based on empirical equation,
for the calculation of the additionally necessary distributor height, ?h3.

Foaming systemIn the case of a foaming system, the foam will be built up in particular in
those areas in which a marked gas injection into the liquid takes place. This applies
particularly to the transfer of the liquid from the feed pipe into the parting box, since
relatively large quantities of liquid are transferred per transition point. Since the description of
the foaming behaviour is very complex, it is advisable to use the foam or system factor, ?,
which is described in the literature.13 This empirical factor is known for numerous mass
transfer tasks and has to be taken into account by Equation 12, based on empirical equation,
for the calculation of the additionally necessary distributor height, ?h3.
ÊåÓû ü
Ù
íÚ
âç
ý
ß
óà
One must notice that system factors, listed in the literature, result from long term experience
in designing tray columns and includes foaming and degassing effects in parallel. Experience
is needed for avoiding overdesigns in taking system factor and degassing into account in
parallel. The increase of liquid distributor height can, however, be markedly restricted if
design methods are taken into account to reduce foaming. For instance, immersed elongated
guide pipes at the feed pipe can be used to feed the liquid into the liquid level of the parting
box and thus reduce foaming. Alternatively, guide sheets can be used as impulse dampers
above the parting box, or a package of structured or random packings can be used within the
parting box or distributor trough in order to support separation of gas and liquid.
Degassing
As has already been described, the high impulse transfer as liquid passes from the feed pipe
into the parting box, causes gas also to be introduced with the jets of liquid into the liquid
layer. The gas then occupies a noticeable volume in the amount of liquid, which causes the
liquid level in the trough to rise. The additional extra height this requires is determined by the
gas portion introduced and by the residence time of the gas in the liquid. The degassing
behaviour is essentially defined by the buoyancy of the gas bubbles, i.e. by the difference in
density between the gas and the liquid. Particularly in high pressure applications the density
differences are small and therefore the degassing efficiency reduced. As is the case with
foaming systems, the additional height, ?h4, which must be taken into account on the basis of
the degassing behaviour can, until now, only be described on the basis of an empirical
equation (Equation 13).
ÊåÓþ ü
Ù
í
Ú
âââ
ÜÝ
ÙÚ ÛÜÝ ÜÞßà
ß
ó
à
The degassing of liquids can, however, be improved by design methods, similar to foaming
behaviour. Wave formationIf the liquid is led from the distributor pipe into the parting box
and then into the distributor troughs, the impulse transfer causes wave formation which is
supported by the flow of the liquid in the troughs. The overall height of a distributor must be
dimensioned so that the wave crests do not lead to a flooding of the distributor troughs or
gas risers. Since the wave formation depends on the quantity of liquid to be distributed, it is
advisable to design the additionally necessary distributor height, ?h5, according to Equation
14 as an empirical function of the liquid load.
ÊåÓÿ üÙÚêÝßà
Taking all the single heights into account, the necessary overall height is defined according to
Equation 15.
Conclusion
Modern liquid distributor designs are relevant for good mass transfer efficiencies in packed
columns. The article describes the flow behaviour of a liquid jet flow that is leaving a
distributor via bottom holes. An equation is provided to describe the coefficient of discharge
and influencing parameters are discussed. The height of a distributor trough has to take into
account the recommended minimum liquid head, specified liquid loading range, gas pressure
drop, foaming and degassing effects and wave creations. This subject is described as well.
ÊåÓÌ ÌïÝ áááááÓîñò åÓì åÓú åÓû åÓþ åÓÿ
Liquid feed of first distributor nozzle dimension

columns. The article describes the flow behaviour of a liquid jet flow that is leaving a
distributor via bottom holes. An equation is provided to describe the coefficient of discharge
and influencing parameters are discussed. The height of a distributor trough has to take into
account the recommended minimum liquid head, specified liquid loading range, gas pressure
drop, foaming and degassing effects and wave creations. This subject is described as well.
Liquid feed of first distributor nozzle dimension
¡¢£¤¤Øì ¥¦§ â¨
©
¡Õò£¤¤Ø âââ
¨
ÐÜ ¢£¤¤Øì
§¦¥¥ ¨
Ï
¡ò£¤¤Ø
ÎÎÎÎÎÎÎÏ
âââ
Ðé Õò£¤¤Ø

¥¦ ¨¨
We accept ¡ò£¤¤Ø ¨¨ Sch 40
Liquid feed second distributor nozzle dimension and outlet nozzle
¡¢£¤¤Øú ¥¦§ â¨
©
¡Õò Ì âââ
¨
ÐÜ ¢£¤¤Øú
§¦¥¥ ¨
Ï
¡ò Ì
ÎÎÎÎÎÎÎÏ
âââ
Ðé Õò Ì

¥¦ ¨¨
We accept ¡ò Ì ¨¨ Sch 40
Gas feed inlet nozzle dimension
¡¢ Ì ç â¨
©
¡Õò! ââ
#$
¢ Ì
é¥¦éç ¨
Ï
¡ò!
ÎÎÎÎÎÎÏ
âââ
Ðé Õò!

%¥¦ç¥§ ¨¨
We accept ¡ò! %¥é¦ ¨¨ Sch 20
Gas feed outet nozzle dimension
¡¢ Ì ç â¨
©
¡ÕòÌ ââ
#$
¢ Ì
é¥¦éç ¨
Ï
¡òÌ
ÎÎÎÎÎÎÏ
âââ
Ðé ÕòÌ

%¥¦ç¥§ ¨¨
We accept ¡ò! %¥é¦ ¨¨ Sch 20

¡Ì ¤! ç%¥¥ ¨¨ Diameter of the tower
¡' Ðç¦ ââ(Ò
¨
Ï Working pressure
¡'ô Ð% ââ(Ò
¨
Ï Design pressure
¡) Ñ 0× Working temperature
¡)ô %¥ 0× Design temperature
ÛÛ1Õ ç§ ¥ 2 Shell material: carbon steel plate
¡3ïÝ çé¥¥ ââ(Ò
¨
Ï Permissible tensile stress
¡ä ÐÐÑ ç¥
4
ââ(Ò
¨
Ï Elastic modulus
¡5 ¥¦ Joint efficiency
¡×Õ % ¨¨ Corrosion allowance
¡6 çé ¨ Overall packing height
¡6ì Ñ¥¥¥ ¨¨ Top disengaging space
¡6ú Ñ¥¥¥ ¨¨ Bottom disengaging space
¡6û Ñ¥¥¥ ¨¨ Middle liquid distribution space
¡7Ì ¤! ááá6 6ì 6ú 6ú Ñ¥ ¨ Shell length
¡689ñ!Ì Ñ¥¥ ¨¨ Skirt length
¡'@ %Ñ¥ ââ(Ò
¨
A
Packing weight
¡6@ì Ñ¥¥¥ ¨¨ First packing section initial distance of applied weight
from the top
¡6@B ¥¥¥ ¨¨ First and second packing height
¡6@ú çç¥¥¥ ¨¨ Second packing section initial distance of applied weight
from the top
¡Cì Ñ¥¥ (Ò Top internals weight
¡6Cì ç¥¥ ¨¨ Top internals distance of applied weight from the top
Middle span internas weight

¡Cú é¥ (Ò Middle span internas weight
¡6Cú ç¥¥¥¥ ¨¨ Middle internals distance of applied weight from the top
¡Cû ç¥ (Ò Bottom span internas weight
¡6Cû ç¥¥¥ ¨¨ Bottom internals distance of applied weight from the top
¡ËDñD¤8EÝïØ¤!8 ç¥ â(Ò
¨
Weight of attachment (pipes, ladders platform)
¡'ì ç%¥ ââ(Ò
¨
Ï Wind pressure up to 20m
¡'ú ç¥ ââ(Ò
¨
Ï Wind pressure beyond 20 m
¡F8Ì ¥¥ ââ(Ò
¨
A
Steel density
¡ËDÝ çÑ (Ò Weight of a plate
6.1 Minimum shell thickness
¡G8 áââ
Ì ¤!
Ñ
Ù
íÚ ÛH
III
PQ
RSTUVW X
ç
ß
óà ×Õ é¦§éç ¨¨
We accept: ¡G8 § ¨¨
Checking maximum pressure allowed before plastic deformation starts to occur
¡ê Yç¦ Minimum elastic deformation alued from the curve
¡'8¤Ý âââââââââââââââ
ÐÐÑ 3ïÝ
ÙÚ ÛG8 ×Õßà
Ù
í
Ú
áç ÐÐç¦ ê
Ù
í
Ú
Ûç ââââ
Ð¥¦Ñ Ì ¤!
7Ì ¤!
ß
ó
à
ß
ó
à
Ðç¥¥ G8
çé¦%çç ââ(Ò
¨
Ï
The alouble pressure is greater than the design pressure, hence the thickness is satisfactory
with respect of plastic deformation
6.2 Torispherical head design
Determine, D, assume values for Rc, Rk and tt
¡`a Ì ¤! ç¦% ¨ Crown radius
¡`9 ¥¦¥§ç% Ì ¤! ¥ ¨¨ Knuckle radius
¡GÌ § ¨¨ Torispherical head thickness
Compute the head Rc/ D, Rk / D, and Rc/ ratios and determine if the following
equations are satisfied.

Compute the head Rc/ D, Rk / D, and Rc/ tt ratios and determine if the following
equations are satisfied.
bb¥¦ ââ
à
Ì ¤!
ç cââ
`9
Ì ¤!
¥¦¥§ bbÑ¥ â
à
GÌ
Ñ¥¥¥
Calculate the geometric constants th, th, R th.d e
¡dÌÍ fghi
Ù
í
Ú
âââââ
ÛÐ¥¦ Ì ¤! `9
Ûà `9
ß
ó
à
ç¦¥
¡eÌÍ âââ
ÎÎÎÎÎÏ Ðà GÌ
`9
ç¦ç¥é
Since , calculate Rth as follows:peÌÍ dÌÍ
¡`ÌÍ ââ
Ì ¤!
Ñ
§¥ ¨¨
Determine the coefficient C1 and C2
¡×ì ÛÐ¦%ç ââ
`9
Ì ¤!
¥¦¥§ ¥¦é
¡×ú ç¦Ñ
Calculate the value of internal pressure expected to produce elastic buckling of the knuckle,Peth
¡'¤ÌÍ ââââââ
ÐÐ×ì ä GÌ
Ï
ÐÐ×ú `ÌÍ
Ù
í
Ú
Ûââ
`ÌÍ
Ñ
`9
ß
ó
à
ç§¦çç ââ(Ò
¨
Ï
Calculate the value of internal pressure that will result in a maximum stress in the knuckle
equal to the material yield strength Py.
Since the allowable stress at design temperature is governed by time-independent properties,
C3 is the material yield strength at the design temperature.
¡×û 3ïÝ
¡'q âââââââ
Ð×û GÌ
ÐÐ×ú `ÌÍ
Ù
í
Ú
Ûââ
`ÌÍ
ÐÑ `9
ç
ß
ó
à
%¦%§ ââ(Ò
¨
Ï
Calculate the value of internal pressure expected to result in a buckling failure of the knuckle,
Pck.
Calculate variable G.
¡r ââ
'¤ÌÍ Ñ¦ç

r
'q

Since , Pck is determenined as:pr ç
¡'a9 Ð
Ù
í
Ú
ââââââââââââââáÛ¥¦¥ r ¥¦Ñ¥%é r
Ï Ð¥¦¥çÑé r
A
áÛáç ¥¦ç¥çé r ¥¦¥%é r
Ï ¥¦¥¥%§ r
A
ß
ó
à
'q §¦çé ââ(Ò
¨
Ï
Calculate the allowable pressure based on a buckling failure of the knuckle, Pak.
¡'ï9 ââ
'a9
ç¦
é¦é% ââ(Ò
¨
Ï
Calculate the allowable pressure based on rupture of the crown, Pac.
¡'ïa ââââ
ÐÐÑ 3ïÝ 5
áâ
`a
GÌ
âç
Ñ
ç¥¦ ââ(Ò
¨
Ï
Calculate the maximum allowable internal pressure, Pa
¡'ï stu ÙÚ v'ï9 'ï9ßà é¦é% ââ(Ò
¨
Ï
'ï é¦é% ââ(Ò
¨
Ï
p'ï 'ô
'ô % ââ(Ò
¨
Ï
The alouble pressure is greater than the design pressure, hence the thickness is satisfactory.
Weight of the torispherical head
¡ËÌ ÐF8Ì
Ù
í
Ú
ÐÐç¦é ââââ
Ð Ì ¤!
Ï
é
GÌ
ß
ó
à
§¦§§ (Ò

6.3 Shell thicnes at difrent heights
Let X be the distance from the top of up to which we can use 6 mm thick shell
1.Circumferential stress induced in shell plate material at a distance X from the top of shell
(due to internal pressure)
¡3wx ââââ
Ð'ô Ì ¤!
Ñ ÙÚ ÛG8 ×Õßà
§¥ ââ(Ò
¨
Ï
y3wx Ð3ïÝ 5
3wx will remain same for entire length
2.Various axial stresses induced due to internal in the shell plate material at a distance X from
the top of the shell.
-Axial stress induced due to internal pressure
¡3€wx ââââ
Ð'ô Ì ¤!
é ÙÚ ÛG8 ×Õßà
%Ñ ââ(Ò
¨
Ï
-Axial stress induced due to dead load
Axial stress induced due to weight of the shell
ÊÊ3€‚ ââââââââââââ
ÐÐâ
é
Ù
Ú ÛÙÚ áÌ ¤! Ñ G8ßà
Ï
Ì ¤!
Ïß
à F8Ì ƒ
â
é
Ù
Ú ÛÙÚ ÛáÌ ¤! Ñ G8 ×Õßà
Ï
Ì ¤!
Ïß
à
ÐF8Ì ƒ
ÊÊ3€‚ ÐF8Ì ƒ ÐÐ¥¦§ ƒ ââ(Ò
¨
Ï
Axial stress induced due to weight of the packing
ÊÊ3€x âââââââââââ
ÐÐÐââââ
Ð Ì ¤!
Ï
é
'@ ƒ ¨
â
é
Ù
Ú ÛÙÚ ÛáÌ ¤! G8 ×Õßà
Ï
Ì ¤!
Ïß
à
Ð§¦Ñ ƒ ââ(Ò
¨
Ï
Axial stress induced due to weight of the internals
Weight of the top internal per surface area
¡Ëñòì ââââ
é Cì
Ð Ì ¤!
Ï ¥¦¥ç ââ(Ò
¨
Ï
Weight of the middle internal per surface area
¡Ëñòú ââââ
é Cú
Ð Ì ¤!
Ï ¥¦¥%é ââ(Ò
¨
Ï

Weight of the bottom internal per surface area
¡Ëñòú ââââ
é Cû
Ð Ì ¤!
Ï ¥¦¥çç ââ(Ò
¨
Ï
Weight of liquid per meter of packing
¢x ç¥¦§¥ â¨
Ó„
Spread of the moving liquid in the packing
¡ËÝñîD âââ
…† ¨
ÐÜ ¢†
ç¦¥%§ ââ¨
A
¨
Weight of liquid per length of the column
ÊÊËÝñîD‡ ÐÐËÝñîD Ü ˆ ÐÐÐç¦¥%% ç¥
A
ˆ â(Ò
¨
Axial stress induced due to weight of liquid
ÊÊÊ3‰E‡ ââââââ
ËÝñîD‡
ÐÐ Ì ¤!
ÙÚ ÛG8 ×Õßà
ââââÐç¥%% ˆ (Ò
çÑ ¨
Ï Ð¦é% ˆ ââ(Ò
¨
Ï
Axial stress induced due to weight of attachments
ÊÊ3‡‰ âââââââ
áËÌ ÐËDñD¤8EÝïØ¤!8 ˆ
ÐÐ Ì ¤!
ÙÚ ÛG8 ×Õßà
Ð‘‘ á¥¦ç ç¦ÑÑé ˆ’’ ââ(Ò
¨
Ï
Total axial stress due to dead loads
Ê3‰ô ááá3‰‡“ 3‰‡†‡ 3‰E‡ 3‡‰
Êáááá‘‘ Ð¥¦§ ˆ’’ ‘‘§¦Ñ ˆ’’ ‘‘¦é% ˆ’’ ‘‘ áç¦ÑÑé ˆ ¥¦ç’’ ‘‘ áá¥¦¥ç ¥¦¥%é ¥¦¥çÑ’’ á¥¦ç ç¦é§ç ˆ
'
ÙÚ Ðç¦ ç¥
” ßà ââ(Ò
¨
Ï
Ê3‰ô á¥¦ç ç¦é§ç ˆ
Axial stress induced due to wind load at a distance X from the top of the shell
ÊÊÊ3‡ âââââââ
ÐÐç¦é ' Ì ¤! ˆ
Ï
ÐÐ Ì ¤!
Ï ÙÚ ÛG8 ×Õßà
ââââââ
Ðç¦é ' ˆ
Ï
ÐÐ Ì ¤!
ÙÚ ÛG8 ×Õßà
Ðç¦é ˆ
Ï

Maximum tensile stress induced in the shell plate material at a distance X from the top of the shell.
ÊÊÊ3Ìîïð Ûá3‰•† 3‡ 3‰ô áÛ%Ñ ‘‘ á¥¦ç ç¦é§ç ˆ’’ Ðç¦é ˆ
Ï áÛ%Ñ ç¦é§ç ˆ Ðç¦é ˆ
Ï
Maximum allowed length for kipping the same shell thickness according to tensile stress.
ÊÊ3Ìîïð Ûá3‰•† 3‡ 3‰ô Ð3ïÝ 5
Ð3ïÝ 5 ÙÚ Ðç¦ç ç¥
A ßà ââ(Ò
¨
Ï ÊÛÛá3‰•† 3‡ 3‰ô Ð3ïÝ 5 ¥
ÊÊÛáÛÐç¦é ˆ
Ï ç¦é§ç ˆ %Ñ çç¥ ÛÛÐç¦é ˆ
Ï ç¦é§ç ˆ § ¥
ÊÊˆ ââââââáÛë ÎÎÎÎÎÎÎÎÏ
ÛëÏ Ðé @
Ñ @
ââââââââââââáç¦é§ç ÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÏ
Ûç¦é§çÏ Ðé ç¦é ‘‘Û§’’
Ñ ç¦é
¡ˆ ââââââââââââáç¦é§ç ÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÏ
Ûç¦é§çÏ Ðé ç¦é ‘‘Û§’’
Ñ ç¦é
%¥¦Ñ
The length x is almost the double off the length of our column hence the same shell
thickness can be used through the column according to tensile stress
Maximum compressive stress induced in the shell plate at a distance X from the top of the
shell.
¡3÷‰ Ðâç
çÑ
ââââä
% ÙÚ Ûç ¥¦%
Ïßà
ââââ
Ì ¤!
Ñç¦ ââ(Ò
¨
Ï Allowable compressive stress
Maximum compressive stress induced in the shell plate material at a distance X from the top
of the shell.
ÊÊÊ3Ìîïð ááÛ3‰•† 3‡ 3‰ô ááÛ%Ñ ‘‘ á¥¦ç ç¦é§ç ˆ’’ Ðç¦é ˆ
Ï ááÛ%Ñ ç¦é§ç ˆ Ðç¦é ˆ
Ï
Maximum allowed length for kipping the same shell thickness according to compressive stress.
ÊÊ3Ìîïð Ûá3‰•† 3‡ 3‰ô Ð3÷‰ 5
ÊÛááÛ3‰•† 3‡ 3‰ô Ð3÷‰ 5 ¥
ÊÊÛÛáÐç¦é ˆ
Ï ç%¦§ç ˆ %Ñ Ñé¥ ÛáÐç¦é ˆ
Ï ç¦é§ç ˆ é§ ¥
ÛëÏ Ðé @
Ñ @
âââââââââââáÛç¦é§ç ÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÏ
Ûç¦é§çÏ Ðé ç¦é é§
Ñ ç¦é
¡ˆ ââââââââââââáÛç¦é§ç ÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÏ
Ûç¦é§çÏ Ðé ç¦é Ûé§
çÑ¦§¥é

Ñ ç¦é
§¥
The length x is less than the length of our column hence we have to use a different thickness
in the half bottom of the column, for this reason we are going to use shell thickness of 8mm in
the half bottom of it and redo the calculation again only for this section of the column.
¡G8 ¨¨
2.Various axial stresses induced due to internal in the shell plate material at a distance X from
the middle of the shell.
-Axial stress induced due to internal pressure
¡3‰•† ââââ
Ð'ô Ì ¤!
é ÙÚ ÛG8 ×Õßà
ç ââ(Ò
¨
Ï
Weight of the top half of the column
¡ËÌ ¤!8Í¤ÝÝ–—ÿ áÐÐâ
é
Ù
Ï
Ì ¤!
Ïß
à ââ
7Ì ¤!
Ñ
F8Ì ËÌ ÙÚ ÐÑ¦§ç ç¥
A ßà (Ò
¡Ë†ï–—ÿ ÐÐ'@ â
é
Ì ¤!
Ï 6@B ÙÚ ÐÑ¦% ç¥
A ßà (Ò
¡ËÌ ¤!8–—ÿ áááËÌ ¤!8Í¤ÝÝ–—ÿ Ë†ï–—ÿ Cì ââ
Cú
Ñ
ÙÚ Ð§¦¥é ç¥
A ßà (Ò
Total axial stress due to weight of half of the column
¡3‡–—ÿ âââââââââââ
ËÌ ¤!8–—ÿ
â
é
Ù
Ú ÛÙÚ ÛáÌ ¤! G8 ×Õßà
Ï
Ì ¤!
Ïß
à
¦ç%é ââ(Ò
¨
Ï
-Axial stress induced due to dead load
Axial stress induced due to weight of the shell
ÊÊ3‰‡“ ââââââââââââ
ÐÐâ
é
Ù
Ï
Ì ¤!
Ïß
à F8Ì ˆ
â
é
Ù
Ú ÛÙÚ ÛáÌ ¤! Ñ G8 ×Õßà
Ï
Ì ¤!
Ïß
à
ÐF8Ì ˆ
ÊÊ3‰‡“ ÐF8Ì ˆ ÐÐ¥¦§ ˆ ââ(Ò
¨
Ï
Axial stress induced due to weight of the packing
ÊÊ3‰‡†‡ âââââââââââ
ÐÐÐââââ
Ð Ì ¤!
Ï
é
'@ ˆ ¨
â Ù
Ú ÛÙÚ ÛáÌ G ×Õßà
Ï
Ì
Ïß
à
Ðé¦çÑ ˆ ââ(Ò
¨
Ï

é ÚÚ áÌ ¤! G8 ×Õà Ì ¤! à
Axial stress induced due to weight of the internals
Weight of the top internal per surface area
¡Ëñòú ââââ
Ñ Cú
Ð Ì ¤!
Ï ¥¦¥ç ââ(Ò
¨
Ï
Weight of the bottom internal per surface area
¡Ëñòú ââââ
é Cû
Ð Ì ¤!
Ï ¥¦¥çç ââ(Ò
¨
Ï
Weight of liquid per meter of packing
¢† ç¥¦§¥ â¨
Ó„
Speed of the moving liquid in the packing
¡ËÝñîD âââ
…† ¨
ÐÜ ¢†
ç¦¥%§ ââ¨
A
¨
Weight of liquid per length of the column
ÊÊËÝñîD‡ ÐÐËÝñîD Ü ˆ ÐÐÐç¦¥%% ç¥
A
ˆ â(Ò
¨
Axial stress induced due to weight of liquid
ÊÊÊ3‰E‡ ââââââ
ËÝñîD‡
ÐÐ Ì ¤!
ÙÚ ÛG8 ×Õßà
ââââÐç¥%% ˆ (Ò
Ñ¥ ¨
Ï Ð¦¥§ç ˆ ââ(Ò
¨
Ï
Axial stress induced due to weight of attachments
ÊÊ3‡‰ âââââââ
áËÌ ÐËDñD¤8EÝïØ¤!8 ˆ
ÐÐ Ì ¤!
ÙÚ ÛG8 ×Õßà
Ð‘‘ á¥¦éÑ§ ¥¦% ˆ’’ ââ(Ò
¨
Ï
Total axial stress due to dead loads
Ê3‰ô ááá3‰‡“ 3‰‡†‡ 3‰E‡ 3‡‰
Êáááá‘‘ Ð¥¦§ ˆ’’ ‘‘é¦çÑ ˆ’’ ‘‘¦¥§ç ˆ’’ ‘‘ á¥¦% ˆ ¥¦éÑ§’’ ‘‘ áá¦ç%é ¥¦¥ç ¥¦¥çç’’ á§¥ ç¥¦Ñ ˆ
Ê3‰ô á§¥ ç¥¦Ñ ˆ

3‰ô á§¥ ¥
Axial stress induced due to wind load at a distance X from the top of the shell
ÊÊÊ3‡ âââââââ
ÐÐç¦é ' Ì ¤! ˆ
Ï
ÐÐ Ì ¤!
Ï ÙÚ ÛG8 ×Õßà
ââââââ
Ðç¦é ' ˆ
Ï
ÐÐ Ì ¤!
ÙÚ ÛG8 ×Õßà
Ð¥¦ç ˆ
Ï
Maximum tensile stress induced in the shell plate material at a distance X from the top of the
shell.
ÊÊÊ3Ìîïð Ûá3‰•† 3‡ 3‰ô áÛç ‘‘ á§¥ ç¥¦Ñ ˆ’’ Ð¥¦ç ˆ
Ï áÛç% ç¥¦Ñ ˆ Ð¥¦ç ˆ
Ï
Maximum allowed length for kipping the same shell thickness according to tensile stress.
ÊÊ3Ìîïð Ûá3‰•† 3‡ 3‰ô Ð3ïÝ 5
Ð3ïÝ 5 ÙÚ Ðç¦ç ç¥
A ßà ââ(Ò
¨
Ï ÊÛÛá3‰•† 3‡ 3‰ô Ð3ïÝ 5 ¥
ÊÊÛáÛÐ¥¦ç ˆ
Ï ç¥¦Ñ ˆ ç% çç¥ ÛÛÐç¦é ˆ
Ï ç%¦§ç ˆ ç¥ ¥
ÛëÏ Ðé @
Ñ @
ââââââââââââáç¥¦Ñ ÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÏ
Ûç¥¦Ñ
Ï Ðé ¥¦ç ‘‘Ûç¥’’
Ñ ¥¦ç
¡ˆ ââââââââââââáç¥¦Ñ ÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÏ
Ûç¥¦Ñ
Ï Ðé ¥¦ç ‘‘Ûç¥’’
Ñ ¥¦ç
éç¦¥ç
The length x is larger than the double off the length of our column hence the same shell
thickness can be used through the column according to tensile stress
Maximum compressive stress induced in the shell plate at a distance X from the top of the
shell.
¡3÷‰ Ðâç
çÑ
ââââä
% ÙÚ Ûç ¥¦%
Ïßà
ââââ
Ì ¤!
é§¦§ç§ ââ(Ò
¨
Ï Allowable compressive stress
Maximum compressive stress induced in the shell plate material at a distance X from the top
of the shell.
ÊÊÊ3Ìîïð ááÛ3‰•† 3‡ 3‰ô ááÛç ‘‘ á§¥ ç¥¦Ñ ˆ’’ Ð¥¦ç ˆ
Ï ááÛç% ç¥¦Ñ ˆ Ð¥¦ç ˆ
Ï
Maximum allowed length for keeping the same shell thickness according to compressive stress.
ÊÊ3Ìîïð Ûá3‰•† 3‡ 3‰ô Ð3÷‰ 5
ÊÛááÛ3‰•† 3‡ 3‰ô Ð3÷‰ 5 ¥
ÊÊÛÛáÐ¥¦ç ˆ
Ï ç¥¦Ñ ˆ ç é§¦§ç§ ÛáÐ¥¦ç ˆ
Ï ç¥¦Ñ ˆ §§ ¥

á¥ ¥ § § § á¥ ¥ §§ ¥
ÛëÏ Ðé @
Ñ @
ââââââââââââáÛç¥¦Ñ ÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÏ
Ûç¥¦Ñ
Ï Ðé ¥¦ç Û§§
Ñ ¥¦ç
¡ˆ ââââââââââââáÛç¥¦Ñ ÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÏ
Ûç¥¦Ñ
Ï Ðé ¥¦ç Û§§
Ñ ¥¦ç
Ñç¦çÑ
ç% ¨C Ñç¦Ñ§ç (¨
Weight of tower
¡ËÌ ¤!8Í¤ÝÝ áÐÐâ
é
Ù
Ï
Ì ¤!
Ïß
à 7Ì ¤! F8Ì ÐÑ ËÌ ÙÚ Ð¦%¥Ñ ç¥
A ßà (Ò
7.0 Design of support
Skirt support
¡89ñ!Ì ç%¥¥ ¨¨ Skirt diameter
¡G89ñ!Ì ¨¨ Skirt thickness
Minimum weight of the vessel with attachments
ÊËÌ ¤!îñò áááËÌ ¤!8Í¤ÝÝ ËDï˜9ñòÞ ËñòÌ ËïÌ˜
Weight of tower
¡G8 § ¨¨
¡ËÌ ¤!8Í¤ÝÝÌ D–—ÿ áÐÐâ
é
Ù
Ï
Ì ¤!
Ïß
à ââ
7Ì ¤!
Ñ
F8Ì ËÌ ÙÚ ÐÑ¦¥¥ ç¥
A ßà (Ò
¡G8 ¨¨
¡ËÌ ¤!8Í¤ÝÝ Ì–—ÿ áÐÐâ
é
Ù
Ï
Ì ¤!
Ïß
à ââ
7Ì ¤!
Ñ
F8Ì ËÌ ÙÚ ÐÑ¦§ç ç¥
A ßà (Ò
¡ËÌ ¤!8Í¤ÝÝ áËÌ ¤!8Í¤ÝÝÌ D–—ÿ ËÌ ¤!8Í¤ÝÝ Ì–—ÿ ÙÚ Ðé¦§ ç¥
A ßà (Ò
Weight of packing
¡ËDï˜9ñòÞ ÐÐÐÑ 6@B ââââ
Ð Ì ¤!
Ï
é
'@ ÙÚ Ð¦é§ ç¥
A ßà (Ò
Weight of internals
¡ËñòÌ ááCì Cú Cû ¥¥ (Ò
Weight of attachments
¡ËïÌ˜ ÐËDñD¤8EÝïØ¤!8 7Ì ¤!
ÙÚ Ð% ç¥
A ßà (Ò
¡ËÌ ¤!îñò áááËÌ ¤!8Í¤ÝÝ ËDï˜9ñòÞ ËñòÌ ËïÌ˜
ÙÚ Ðç¦éé ç¥
” ßà (Ò
Minimum weight of the vessel with attachments

ËÌ ¤!îñò áááËÌ ¤!8Í¤ÝÝ ËDï˜9ñòÞ ËñòÌ ËïÌ˜ Ú ¥ à Ò
Maximum weight of the vessel
Weight of water during testing
¡ÜïÌ¤! ç¥¥¥ ââ(Ò
¨
A
¡ËÝñ™ ÐÐÜïÌ¤! ââââ
Ð Ì ¤!
Ï
é
7Ì ¤!
ÙÚ ÐÑ¦§ ç¥
” ßà (Ò
¡ËÌ ¤!îïð ááËÌ ¤!8Í¤ÝÝ ËïÌ˜ ËÝñ™
ÙÚ Ð%¦éÑ ç¥
” ßà (Ò
Applied axial force
¡ø‰ ÐÛËÌ ¤!îïð Ò ÐÛ%¦%é ç¥
d
2
Applied net section bending moment
Wind loads acting over the vessel
¡eì ç
¡eú ¥¦ ¡øì ÐÐÐeì eú 'ì ÙÚ áÌ ¤! Ñ G8ßà çç¦§ â(Ò
¨
¡øú ÐÐÐeì eú 'ú ÙÚ áÌ ¤! Ñ G8ßà ç%¦ç â(Ò
¨
Applied wing bending moment
¡fì ÐÐøì 7Ì ¤! ââ
7Ì ¤!
Ñ
ÙÚ ÐÑ¦% ç¥
” ßà Ð(Ò ¨
¡fú ÐÐøú 689ñ!Ì
Ù
í
Ú
á7Ì ¤! ââ
689ñ!Ì
Ñ
ß
ó
à
ÙÚ Ð¦%éç ç¥
A ßà Ð(Ò ¨
¡f áfì fú ÙÚ Ð%¦çÑ ç¥
” ßà Ð(Ò ¨
Pressure loads
¡'89ñ!Ì ¥
Determine applicability of the rules of VIII-2, based on satisfaction of the following
requirements.
The section of interest is at least 2.5 Rt away from any major structural discontinuity.
¡g89ñ!Ì ââ
89ñ!Ì
Ñ ¡h ÐÑ¦ ÎÎÎÎÎÎÎÎÎÏ Ðg89ñ!Ì G89ñ!Ì ¥¦ç ¨
Shear force is not applicable.
The shell R / t ratio is greater than 3.0:
g

iÊââ
g89ñ!Ì
G89ñ!Ì
§ %
Calculate the membrane stress for the skirt, with a weld joint efficiency of E =0.85.
Note that the maximum bending stress occurs at .Êj ¥ èHÒ
¡j ¥ èHÒ ¡3kî Ðâââââââ
'89ñ!Ì
Ð5 lm
Ù
í
Ú
âââââ
á89ñ!Ì ÐÑ G89ñ!Ì
89ñ!Ì
ß
ó
à
ââ(Ò
¨
Ï ¥ ââ(Ò
¨
Ï
¡ 89ñ!Ì á89ñ!Ì ÐÑ G89ñ!Ì
¡38îì âç
5
Ù
í
íÚ
ááâââââââ
ÐÐ'89ñ!Ì 89ñ!Ì (Ò
Ð¨ ÙÚ Û 89ñ!Ì
Ï 89ñ!Ì
Ïßà
âââââââ
Ðé ËÌ ¤!îïð
Ð ÙÚ Û 89ñ!Ì
Ï 89ñ!Ì
Ïßà
ââââââââ
ÐÐÐ%Ñ f 89ñ!Ì nop ‘‘j’’
Ð ÙÚ Û 89ñ!Ì
”
89ñ!Ì
” ßà
ß
ó
óà
é§§¦§ ââ(Ò
¨
Ï
¡38îú âç
5
Ù
í
íÚ
Ûáâââââââ
ÐÐ'89ñ!Ì 89ñ!Ì (Ò
Ð¨ ÙÚ Û 89ñ!Ì
Ï 89ñ!Ì
Ïßà
âââââââ
Ðé ËÌ ¤!îïð
Ð ÙÚ Û 89ñ!Ì
Ï 89ñ!Ì
Ïßà
ââââââââ
ÐÐÐ%Ñ f 89ñ!Ì nop ‘‘j’’
Ð ÙÚ Û 89ñ!Ì
”
89ñ!Ì
” ßà
ß
ó
óà
ÛÑÑÑ¦ç% ââ(Ò
¨
Ï
¡fÌ ¥
¡q ââââââââââ
ÐÐÐÐÐç§ fÌ 89ñ!Ì nop ‘‘j’’ 7 ââ
(Ò
¨
Ï
Ð ÙÚ Û 89ñ!Ì
”
89ñ!Ì
” ßà
¥ ââ(Ò
¨
Ï
Calculate the principal stresses.
¡3ì ¥¦
Ù
íÚ áá3kî 38îì
ÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÏ
áÙÚ Û3kî 38îìßà
Ï
Ðé q
Ï ß
óà é§§¦§ ââ(Ò
¨
Ï
¡3ú ¥¦
Ù
íÚ Ûá3kî 38îì
ÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÏ
áÙÚ Û3kî 38îìßà
Ï
Ðé q
Ï ß
óà ¥ ââ(Ò
¨
Ï
¡3û ÐÐ¥¦ '89ñ!Ì ââ(Ò
¨
Ï ¥ ââ(Ò
¨
Ï
Check the allowable stress acceptance criteria.
¡3r ââç
ÎÎÏ
Ñ
ÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÎÏ Ù
Ú ááÙÚ Û3ì 3úßà
Ï
ÙÚ Û3ú 3ûßà
Ï
ÙÚ Û3û 3ìßà
Ïß
à é§§¦§ ââ(Ò
¨
Ï
s3r 3ïÝ
Note that VIII-2 uses an acceptance criteria based on von Mises Stress. VIII-1 typical uses the
maximum principal stress in the acceptance criteria. Therefore:
tuv ÙÚ ww3ì 3ú 3ûßà é§§¦§ ââ(Ò
¨
Ï
(

stuv ÙÚ ww3ì 3ú 3ûßà 3ïÝ 3ïÝ
ÙÚ Ðç¦é ç¥
A ßà ââ(Ò
¨
Ï
For cylindrical and conical shells, if the axial membrane stress, is compressive, then38îì
VIII-2, Equation (4.3.45) shall be satisfied where Fxa is evaluated using paragraph 4.4.12.2
with .¡æ ¥¦ç
x38îì øðï
For that is compressive, a buckling check is required.38îú
VIII-2, paragraph 4.4.12.2.b – Axial Compressive Stress Acting Alone.
In accordance with VIII-2, paragraph 4.4.12.2.b, the value of is calculated as follows, withøðï
.¡æ ¥¦ç
The design factor FS used in VIII-2, paragraph 4.4.12.2.b is dependent on the predicted
buckling stress and the material’s yield strength, as shown in VIII-2, paragraph 4.4.2.øñ˜ 3ïÝ
An initial calculation is required to determine the value of by setting FS =1.0 , withøðï
= . The initial value of is then compared to as shown in paragraph 4.4.2 andøñ˜ øðï øñ˜ 3ïÝ
the value of FS is determined. This computed value of FS is then used in paragraph
4.4.12.2.b. For ¡æ ¥¦ç
Êøðï tym ÙÚ wøðïì øðïúßà
âââ
89ñ!Ì
G89ñ!Ì
ç§é¦
¡g 89ñ!Ì âââ
89ñ!Ì
Ñ
¡fð âââââ
Ñ 689ñ!Ì
ÎÎÎÎÎÎÎÎÎÏ Ðg 89ñ!Ì G89ñ!Ì
§¦ç
Since , calculate xa1 F as follows with an initial value of FS=1.xxç% âââ
89ñ!Ì
G89ñ!Ì
§¥¥
¡ø© ç
¡øðïì ââââââ
é§§ 3ïÝ
Ðø©
Ù
í
Ú
á%%ç âââ
89ñ!Ì
G89ñ!Ì
ß
ó
à
ÙÚ Ðç¦%ç ç¥
A ßà ââ(Ò
¨
Ï
The value of is calculated as follows with an initial value of FS =1.øðïú
Êøðïú ââââ
ÐÐ×ð ä G89ñ!Ì
Ðø© 89ñ!Ì
Since , calculate C x as follows:xâââ
89ñ!Ì
G89ñ!Ì
çÑé
Since , calculate c. as follows:ifð ç
¡×ð âââââÐé¥ z
¥¦%¡z ç

×ð
á% âââ
89ñ!Ì
G89ñ!Ì
¥ %
Therefore:
¡øðïú ââââ
Ðø© 89ñ!Ì
ÙÚ Ð¦é ç¥
A ßà ââ(Ò
¨
Ï
¡øðï tym ÙÚ wøðïì øðïúßà ÙÚ Ðç¦%ç ç¥
A ßà ââ(Ò
¨
Ï
With a value of , in accordance with VIII-2, paragraph 4.4.2, the value of FS isÊøñ˜ øðï
determined as follows.
¡øñ˜ çç¦Ñ ââ(Ò
¨
Ï Ð¥¦ 3ïÝ ¥ ââ(Ò
¨
Ï
s3ïÝ øñ˜
¡ø1 ÛÑ¦é¥ Ð¥¦éç
Ù
í
Ú
ââ
øñ˜
3ïÝ
ß
ó
à
Ñ¦é¥ç
Using this computed value of FS =2.401 in paragraph 4.4.12.2.b, is calculated as follows.øðï
¡øðïì ââââââ
Ðé§§ 3ïÝ
Ðø1
Ù
í
Ú
á%%ç âââ
89ñ!Ì
G89ñ!Ì
ß
ó
à
é¦éé ââ(Ò
¨
Ï
¡øðïú ââââ
Ðø1 89ñ!Ì
ÙÚ Ð%¦éÑ ç¥
A ßà ââ(Ò
¨
Ï
¡øðï tym ÙÚ wøðïì øðïúßà é¦éé ââ(Ò
¨
Ï
Compare the calculated axial compressive membrane stress, to the allowable axial38îì
compressive membrane stress, per following criteria:øðï
x38îì øðï
Therefore, local buckling due to axial compressive membrane stress is not a concern.

7.0' Design of support (conventional method)
Periods of vibration at minimum weight
¡)îñò ââââââââââââç
ÐÐÐ§¦% ç¥
{d
Ù
í
Ú
ââ
7Ì ¤!
Ì ¤!
ß
ó
à
|}d ÎÎÎÎÎÎÎÎÎÎÏ
ââââ
ÐËÌ ¤!îñò Ò
ÐG8 (Ò
¥¦¥§Ñ ©
x)îñò ¥¦ © Hence ¡eD ç
Periods of vibration at maximum weight
¡)îñò ââââââââââââç
ÐÐÐ§¦% ç¥
{d
Ù
í
Ú
ââ
7Ì ¤!
Ì ¤!
ß
ó
à
|}d ÎÎÎÎÎÎÎÎÎÎÏ
ââââ
ÐËÌ ¤!îïð Ò
ÐG8 (Ò
¥¦¥é ©
x)îñò ¥¦ © Hence ¡eD ç
Minimum skirt thickness
stress due to wind load
¡( ¥¦ Coefficient of wind influence for cylindrical surface
For minimum weight condition
¡'îñò ÐÐÐÐ( eD 'ì Ì ¤! 7Ì ¤!
ÙÚ ÐÑ¦%§§ ç¥
A ßà (Ò
For maximum weight condition
¡'îïð ÐÐÐÐ( eD 'ì ÙÚ áÌ ¤! ÐÑ ¥¦ç ¨ßà 7Ì ¤!
ÙÚ ÐÑ¦% ç¥
A ßà (Ò
Minimum wind moment
¡fîñò Ð'îñò
Ù
í
Ú
áââ
7Ì ¤!
Ñ
689ñ!Ì
ß
ó
à
ÙÚ ÐÑ¦ ç¥
” ßà Ð(Ò ¨
Maximum wind moment
¡fîïð Ð'îïð
Ù
í
Ú
áââ
7Ì ¤!
Ñ
689ñ!Ì
ß
ó
à
ÙÚ Ð%¦éç% ç¥
” ßà Ð(Ò ¨
Determination of skirt thickness
Minimum and maximum stress on the skirt
Minimum and maximum wind load stress on the skirt
Assuming a small skirt thickness Din=Dout we can write:
Minimum wind load stress ÊÊ3‡îñò âââââ
Ðé fîñò
ÐÐ Ì ¤!
Ï G89ñ!Ì
ââÑÑÑ
G89ñ!Ì
ââ(Ò
¨
Maximum wind load stress ÊÊ3‡îïð âââââ
Ðé fîïð
ÐÐ Ì ¤!
Ï G89ñ!Ì
ââÑç
G89ñ!Ì
ââ(Ò
¨
Minimum and maximum dead load stress on the skirt

Minimum and maximum dead load stress on the skirt
Minimum dead load stress ÊÊ3ôîñò âââââ
ËÌ ¤!îñò
ÐÐ Ì ¤! G89ñ!Ì
ââ%¦Ñ
G89ñ!Ì
ââ(Ò
¨
Maximum dead load stress ÊÊ3ôîïð âââââ
ËÌ ¤!îïð
ÐÐ Ì ¤! G89ñ!Ì
âââ%¦ç
G89ñ!Ì
ââ(Ò
¨
Minimum and maximum tensile stress without any eccentric load
Maximum tensile stress ÊÊ3“îïð Û3‡îñò 3ôîñò ââÑç%
G89ñ!Ì
ââ(Ò
¨
Minimum tensile stress ÊÊ3“îñò Û3‡îïð 3ôîïð ââÑé
G89ñ!Ì
ââ(Ò
¨
Determination of skirt thickness according to minimum and maximum tensile stress
Taking into account the joint efficiency of 0.7 we can derive:
¡5 ¥¦
¡3“ïÝ Ð3ïÝ 5 ÙÚ Ðç¦ç ç¥
A ßà ââ(Ò
¨
Ï
Also we can write
ÊÊ3“ïÝ âââ
3“îïð
G89ñ!Ì
ââÑé
G89ñ!Ì
ââ(Ò
¨
ÊÊÊG89ñ!Ì âââ
3“îïð
3“ïÝ
ââÑé
çç¥
¨ Ñ¦¥ ¨¨
Minimum and maximum stress due to compressive loads
Minimum compressive stress ÊÊ3÷“îñò á3‡îñò 3ôîñò ââÑÑ§%
G89ñ!Ì
ââ(Ò
¨
Maximum compressive stress ÊÊ3÷“îïð á3‡îïð 3ôîïð ââÑ§
G89ñ!Ì
ââ(Ò
¨
Determination of skirt thickness according to minimum and maximum compressive stress
We can write:
ÊÊ3÷“ âââââ
ÐÐ¥¦çÑ ä G89ñ!Ì
Ì ¤!
ââÑÑ§%
G89ñ!Ì
ââ(Ò
¨
From this we can derive ÊG89ñ!Ì
Ï ââââ
ÐÌ ¤! ÑÑ§%
Ð¥¦çÑ ä
ââ(Ò
¨
Î

¡G89ñ!Ìîñò
ÎÎÎÎÎÎÎÎÎÎÎÎÎÎÏ
ââââ
ÐÌ ¤! ÑÑ§%
Ð¥¦Ñ ä
ââ(Ò
¨
¦§ç ¨¨
¡G89ñ!Ìîïð
ÎÎÎÎÎÎÎÎÎÎÎÎÎÎÏ
ââââ
ÐÌ ¤! Ñ§
Ð¥¦Ñ ä
ââ(Ò
¨
¦%¥ ¨¨
We accept skirt thickness: ¡G89ñ!Ì ç¥ ¨¨ Taking into consideration corrosion and other
construction factors
Design of skirt bearing plates
Maximum compressive stress between bearing plate and foundation
¡89ñ!Ì áÌ ¤! Ñ G89ñ!Ì ÙÚ Ðç¦%Ñ ç¥
A ßà ¨¨
¡1 Ñ¥¥ ¨¨ Distance between outer radius of bearing plate and inner radius of bearing plate
Ê3÷“~† áââââ
ËÌ ¤!îïð
Õ
âââ
fîïð

Were:
¡Õ âââââââââ
Ð
Ù
Ú ÛÙÚ á89ñ!Ì Ñ 1ßà
Ï
89ñ!Ì
Ïß
à
é
ÙÚ Ð¦ ç¥
A ßà ¨
Ï
¡ â
%Ñ
ââââââââ
Ù
Ú ÛÙÚ á89ñ!Ì Ñ 1ßà
”
89ñ!Ì
” ß
à
á89ñ!Ì Ñ 1
ÙÚ Ð%¦Ñ§% ç¥
d ßà ¨
A
From this we can derive:
¡3÷“~† áââââ
ËÌ ¤!îïð
Õ
âââ
fîïð

çé¦¥éç ââ(Ò
¨
Ï
x3÷“~† Ð% ââ(Ò
¨
Ï Which is concrete compression stress allowed
¡ D á89ñ!Ì Ñ 1 ÙÚ Ðç¦Ñ ç¥
A ßà ¨¨ Bearing plate outer diameter
Bearing plate thickness
¡3ïÝ Ðç ââ(Ò
¨
Ï Allowable compressive stress for bearing plate
¡G D
ÎÎÎÎÎÎÎÎÎÎÏ
ââââ
Ð% 3÷“~† 1
Ï
3ïÝ
%Ñ¦¥ ¨¨
We accept: ¡G D é¥ ¨¨
As bearing thickness plate is less/equal than 20 mm gaskets are no required.
Minimum compressive stress in bearing plate

Required: “CONFIDENTIAL” Sh.a. DESIGNED: Arberor MITA Design group“ELMAGERARD”As bearing thickness plate is less/equal than 20 mm gaskets are no required.
Minimum compressive stress in bearing plate
¡3÷“~†îñò áââââ
ËÌ ¤!îñò
Õ
âââ
fîïð

çç¦§ ââ(Ò
¨
Ï
Anchorage factor of overturn
¡5 âââââââ
ÐÐËÌ ¤!îñò ¥¦éÑ D
fîñò
¥¦%Ñ
s5 ç¦ This means that akore bolts are required
Calculation of anchor bolts
Anchor bolts sum of actual force
¡2ï Ð3÷“~†îñò Õ ÙÚ Ðç¦çé% ç¥
d ßà (Ò
¡3÷“~ é¥ ââ(Ò
¨
Ï Hot rolled plain carbon steel allowable stress
The sum area of bolts
¡Õ8 ââ
2ï
3÷“~
ÙÚ ÐÑ¦çç ç¥
” ßà ¨¨
Ï
Determine bolt diameter
Maximum allowable number of bolts is 20, from this we can derive the surface of a bolt
¡Ö Ñ¥
¡Õ ââ
Õ8
Ö
ÙÚ Ðç¦¥ ç¥
A ßà ¨¨
Ï
¡
ÎÎÎÎÎÏ
ââ
é Õ

%§¦¥ ¨¨
We accept M36 for anchor bolt
Thickness of bearing plate inside the bolting chair
ÊÊG ˜ ââââ
ÎÎÎÎÎÎÎÏ Ð§ fîïð
ÎÎÎÎÎÎÏ Ð1 3ïÝ
ââââ
ÎÎÎÎÎÎÎÎÏ
Ð§ âââ
Ð2ï ø

¡ø ç¥ ¨¨ Spacing between stiffeners
¡G ˜ ââââââ
ÎÎÎÎÎÎÎÎÎÎÎÎÏ
Ð§ ââââ
ÐÐ2ï ø ¨¨

Ñ¦Ñ§ ¨¨
We accept: ¡G ˜ %¥ ¨¨

8.0 Design of flanged joints
General data
¡'ô Ð% ââ(Ò
¨
Ï Design pressure
ÛÛ1Õ ç§ ¥ 2 Shell material: carbon steel plate
¡3ïÝ 8¤Ý çé¥¥ ââ(Ò
¨
¡ä ÐÐÑ ç¥
4
ââ(Ò
¨
Ï Elastic modulus
¡5 ¥¦ Joint efficiency
Flange data
ÛÛ1Õ ç¥ Ñ¥¥ Shell material: carbon steel plate
¡3ïÝ £Ýï çé¥¥ ââ(Ò
¨
¡ä ÐÐÑ ç¥
4
ââ(Ò
¨
Ï Elastic modulus
Bolt data
ÛÛ1Õ ç% € Ñ¥¥ Shell material: carbon steel plate
¡3ïÝ ÝÌ ç¥ ââ(Ò
¨
¡ä ÐÐÑ ç¥
4
ââ(Ò
¨
Ï Elastic modulus
¡ Ñ¥ ¨¨ Diameter of the bolt
¡2 %¥ Number of bolts
¡Õ! ç ¨¨
Ï Root Area Of the bolt
Gasket data
ü‚@G¨HG@‚ƒ@(HGHè Shell material: iron/soft steel
¡¨ %¦ Gasket factor
¡„ Ð% ââ(Ò
¨
Ï Seating stress
¡Þñ ç%Ñ¥ ¨¨ Inside diameter

Þñ % ¥ Inside diameter
¡Þ ç%¥ ¨¨ Outside diameter
Design rules for bolted flange connections with ring type gaskets are provided in VIII-1
Mandatory Appendix 2. The rules in this paragraph are the same as those provided in VIII-2.
The design procedures in VIII-2, paragraph 4.16 are used in this example problem with
substitute references made to VIII-1 Mandatory Appendix 2 paragraphs.
Evaluate the girth flange in accordance with VIII-2, paragraph 4.16.
VIII-2, paragraph 4.16.6, Design Bolt Loads. The procedure to determine the bolt loads for
the operating and gasket seating conditions is shown below.
Determine the design pressure and temperature of the flanged joint.
¡'ô Ð% ââ(Ò
¨
Ï Design preasure
¡)ô Ñ 0× Design temperature
Select a gasket and determine the gasket factors m and y from Table 4.16.1 (VIII-1,Table
2-5.1).
¡¨ %¦ Gasket factor
¡„ Ð% ââ(Ò
¨
Ï Seating stress
Determine the width of the gasket, N , basic gasket seating width, bo , the effective
gasket seating width, b , and the location of the gasket reaction, G
¡2 âââ
ÛÞ Þñ
Ñ
ç ¨¨
From Table 4.16.3 (VIII-1, Table 2-5.2), Facing Sketch Detail 2, Column II,
¡ã£ % ¨¨ raised nubbin width
¡ë– âââ
áã£ Ð% 2

§ ¨¨
For së– §¦% ¨¨ ¡ë ë–
Therefore, the location of the gasket reaction is calculated as follows (VIII-1, paragraph 2-3).
G mean diameter of the gasket contact face
¡… âââ
áÞ Þñ
Ñ
ÙÚ Ðç¦%% ç¥
A ßà ¨¨
Determine the design bolt load for the operating condition, (VIII-1, paragraph 2-5).
¡Ë– áÐÐâ
…
Ï 'ô ÐÐÐÐ … 'ô ë ¨ ÙÚ Ðé¦éÑ ç¥
” ßà (Ò

Ë– á
é
… ô … ô ë ¨ Ú ¥ à Ò
Determine the design bolt load for the gasket seating condition (VIII-1, paragraph 2-5).
ÊËÞ
Ù
í
Ú
âââ
áÕî Õ
Ñ
ß
ó
à
1 Þ
Were:
¡Õ ÐÕ! 2 ¦ ¨
Ï
ÊÕî tuv
Ù
í
í
íÚ
w
Ù
í
í
íÚ
ââââââ
ááË– ø‰ ââ
é f¤
…
3ïÝ ÝÌ
ß
ó
ó
óà
Ù
í
Ú
âââ
ËÞ8
3ïÝ ÝÌ
ß
ó
à
ß
ó
ó
óà
Axial load aplued in flange joint
¡ø‰ ÛËÌ ¤!îïð
Ù
íÚ
áÐÐâ
é
Ù
Ï
Ì ¤!
Ïß
à 6û F8Ì ËÌ
ß
óà
ÙÚ Ð%¦%§ ç¥
” ßà (Ò
Bending moment aplued in flange joint
¡fì ÛÐøì
Ù
í
Ú
ÐÙÚ Û7Ì ¤! 6ûßà ââââ
Û7Ì ¤! 6û
Ñ
ß
ó
à
ÐÐøì 6û ââ
6û
Ñ
ÙÚ Ðç¦ç§ ç¥
” ßà Ð(Ò ¨
¡fú ÐÐøú 689ñ!Ì
Ù
í
Ú
áÛ7Ì ¤! 6û âââââââ
ÙÚ ÛÛ7Ì ¤! 6û 689ñ!Ìßà
Ñ
ß
ó
à
ÙÚ Ð¦ ç¥
A ßà Ð(Ò ¨
¡f¤ áfì fú ÙÚ ÐÑ¦¥§ ç¥
” ßà Ð(Ò ¨
Maximum load aplued in flange joint
¡ËÞ8 ÐÐÐ ë … „ ÙÚ Ðç¦%é§ ç¥
d ßà (Ò
From this we can derive
¡Õî tuv
Ù
í
í
íÚ
w
Ù
í
í
íÚ
ââââââ
ááË– ø‰ ââ
é f¤
…
3ïÝ ÝÌ
ß
ó
ó
óà
Ù
í
Ú
âââ
ËÞ8
3ïÝ ÝÌ
ß
ó
à
ß
ó
ó
óà
Ñé¦%¥Ñ ¨¨
Ï
Bolt load for the operating condition
¡ËÞ
Ù
íâââ
áÕî Õ ß
ó 3ïÝ ÝÌ ÙÚ Ð¦ ç¥
d ßà (Ò

ËÞ í
Ú Ñ
ó
à
3ïÝ ÝÌ Ú ¥ à Ò
Flange Design Procedure. The procedure in this paragraph can be used to design circular
integral, loose or reverse flanges, subject to internal or external pressure, and external
loadings. The procedure incorporates both a strength check and a rigidity check for flange
rotation.
Determine the design pressure and temperature of the flanged joint and the external
net-section axial force, Fa , and bending moment, Me.
¡'ô Ð% ââ(Ò
¨
Ï Design pressure
ø‰
ÙÚ Ð%¦%§ ç¥
” ßà (Ò Axial force
f¤
ÙÚ ÐÑ¦¥§ ç¥
” ßà Ð(Ò ¨ Bending moment
Determine the design bolt loads for operating condition , and the gasket seating conditionË–
, and the corresponding actual bolt load area , (VIII-1, paragraph 2-5).ËÞ Õ
Ë–
ÙÚ Ðé¦éÑ ç¥
” ßà (Ò
ËÞ ÙÚ Ð¦ ç¥
d ßà (Ò
Õ ¥¦¥¥§ ¨
Ï
Determine an initial flange geometry (see Figure E4.16.1) in addition to the information
required to determine the bolt load, the following geometric parameters are required, (VIII-1,
paragraph 2-3).
¡€ áç%ç§ ¨¨ Ñ % ¨¨ ç¦%ÑÑ ¨ Flange bore
¡× ç¥¥ ¨¨ Bolt circle diameter
¡Õ ç§¥¥ ¨¨ Flange outside diameter
¡G ¥ ¨¨ Flange thicness
¡Òì ÛÐ¥¦ ‘‘ Ûç%¥ ¨¨ ç%ç§ ¨¨’’ ×Õ çé ¨¨ Thickness of the hub at the large end
¡Ò– Ûç¥ ¨¨ ×Õ ¨¨ Thickness of the hub at the small end

Ò– ¥ × Thickness of the hub at the small end
¡Ó ¨¨ Hub length
Determine the flange stress factors using the equations in Tables 4.16.4 and 4.16.5 (VIII-1,
Table 2-7.1 and Fig. 2-7.1 – Fig. 2-7.6).
¡e âÕ
€
ç¦Ñç
¡† ââç
Ûe ç
Ù
í
Ú
á¥¦§§é Ðç¦ç§¥ ââââÐe
Ï lo‡ ‘‘e’’
Ûe ç
ß
ó
à
¦%
¡) ââââââââââÛÐe
Ï ‘‘ áç Ð¦Ñé§ lo‡ ‘‘e’’’’ ç
ÐÙÚ áç¦¥éÑ Ðç¦éé e
Ïßà ‘‘ Ûe ç’’
ç¦%
¡ê ââââââââââÛÐe
Ï ‘‘ áç Ð§¦ÑÑé§ lo‡ ‘‘e’’’’ ç
ÐÐç¦%§ç%§ ÙÚ Ûe
Ï çßà ‘‘ Ûe ç’’
¦ç
¡ âââáe
Ï ç
Ûe
Ï ç
¦%¥%
VIII-1, Fig. 2-7.2:
¡Ó– ÎÎÎÎÏ Ð€ Ò– §¦ç ¨¨
¡ˆÞ â
Òì
Ò–
Ñ
¡ˆÍ âÓ
Ó–
¥¦Ñ
¡øB áááÛ¥¦§ ¥¦Ñ¥çÑ lm ÙÚˆÞßà ÐÐ¦Ñ ç¥
{A
lm ÙÚˆÍßà Ð¥¦çÑ%§ ÙÚlm ÙÚˆÞßàßà
Ï
Ð¥¦¥% ÙÚlm ÙÚˆÍßàßà
Ï
¡øBB áÛÐÐÛ¥¦çééÑ lm ÙÚˆÞßà lm ÙÚˆÍßà Ð¥¦¥ççÑ ÙÚlm ÙÚˆÞßàßà
A
Ð¥¦¥çÑ%§ ÙÚlm ÙÚˆÍßàßà
A
¡øBBB áÐÛ¥¦¥%§% lm ÙÚˆÞßà ÙÚlm ÙÚˆÍßàßà
Ï
ÐÐ¥¦¥Ñ%ç ÙÚlm ÙÚˆÞßàßà
4
lm ÙÚˆÍßà
¡ø ááøB øBB øBBB ¥¦çé
For xx¥¦ ˆÍ Ñ
¡ÔB ááááÛÛ¥¦¥çéé§ âââ¥¦ç%
ˆÞ
ââââ¥¦¥é§çç
ˆÍ
âââ¥¦§¥ç
ˆÞ
Ï ââââ¥¦¥ÑÑ
ˆÍ
Ï âââ¥¦Ñéé%ç%
ÐˆÞ ˆÍ
âââ¥¦çç%Ñ
ˆÞ
A
¡ÔBB ÛÛÛââââ¥¦¥¥ÑÑ§
ˆÍ
A
ââââ¥¦¥Ñ§§Ñ%
ÐˆÞ ˆÍ
Ï âââ¥¦Ñç¥¥
ÐˆÞ
Ï ˆÍ

Þ Þ
¡Ô áÔB ÔBB ¥¦Ñçç
¡ü tuv
Ù
í
íÚ
wç ââââââââââââââââââââââââââââ
ÛáááÛ¥¦¥Ñ Ð¥¦¥%%§§%% ˆÞ Ð¥¦§éç§ ˆÞ
Ï Ð¥¦¥§§Ñ§ ˆÍ Ð¥¦%é¥é ˆÍ
Ï Ðé¦ç§ ˆÍ
A
áááÛç ÐÐ¦§¥% ç¥
{A
ˆÞ Ðç¦§Ñ¥é ˆÍ Ð%¦é%Ñ ˆÍ
Ï Ðç¦%¥Ñ ˆÍ
A
ß
ó
óà
ü ç
VIII-1, paragraph 2-3:
¡è âââââ
ÐÐê
Ù
í
Ú
â
Ò–
CÖ
ß
ó
à
Ï
Ó–
Ô
çÑ¦§ CÖ
¡H ââø
â
Ó–
CÖ
¥¦Ñç
¡7 âââ
áÐâG
CÖ
H ç
)
âââ
Ù
íÚ
âG
CÖ
ß
óà
A
âè
CÖ
Ñ¦ÑÑ
Determine the flange forces, (VIII-1, paragraph 2-3).
¡6ô ââââ
ÐÐ €
Ï 'ô
é
ÙÚ Ðé¦çç ç¥
” ßà (Ò
¡6 ââââ
ÐÐ …
Ï 'ô
é
ÙÚ Ðé¦ç ç¥
” ßà (Ò
¡6 Û6 6ô ç%¦% (Ò
¡6$ ÛË– 6ô ÙÚ Ð%¦§é ç¥
A ßà (Ò
Determine the flange moment for the operating condition using Equation (4.16.14) or
Equation (4.16.15), as applicable (VIII-1, paragraph 2-6). In these equations, , , , areÓô Ó Ó$
determined from Table 4.16.6 (VIII-1, Table 2-6). For VIII-2 designs – For integral and loose
type flanges, the moment, Moe is calculated using Equation (4.16.16) where I and Ip, in this
equation are determined from Table 4.16.7.
Êf ¤ áÐÐÐé f¤
Ù
íââââˆ ß
ó
Ù
íâââ
Óô ß
ó Ðø‰ Óô

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