Production of 20 kton/a TMP (trimethylolpropane) from n-butyraldehyde and formaldehyde

(1)

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FVO Nr.

1

3191

1

"F abrieksvoorontwerp"

Department of Chemical Process Technology

Subject

Production of20 ktonJa IMP (trimethylolpropane) from

n-butyraldehyde and formaldehyde.

Autors

C. (Kees) van de Beek

J

.

(Jurg) Bremmer

Phone number

Il

J. W. (Jillis) Raadschelders

L

M. (Martin) Reuvers

0172-215895

.

015-2621135

023-5371136

0180-411055

Keywords

TMP, 2-ethyl-2-hydroxymethyl-propane-l,3-diol,

Polyhydric alcohols, Polyurethanes, Solvent extraction

Start Project :

Date Report :

11 October 1996

(2)

.~

Errata and Supplements - FV 0 3191

page i, line 4: The reaction section consists of three reactors in serie .. .

replace by: The reaction section consists of three reactors in series .. .

page ii: 6.1. Toxiciy of the Chemicals

replace by: Toxicity of the Chemicals

page 1-1: In fig. 1, the O-atom in the carboxyl group is missing.

page 1-2, line 4: Gypsum and formic acid are formed (see reaction 8).

replace by: Gypsum and formic acid are formed (see reaction 8, page 2-3).

page 2-4, table 3: the price of CaS04 has been estimated at 0.2 $/kg

page 3-1, paragraph 2, line 8: ... with a cascade of three destilation towers.

replace by: ... with a cascade of three extraction towers.

page 3-2, line 1: ... resulted in (only) two artic1es, ...

supplement: ... resulted in (only) two artic1es [lit. 11

+

12].

page 3-2, chapter 3.3, line 5: The units ca1culated with MS Excel ...

replace by: The other units ca1culated with MS Excel ...

page 4-2, under equation 4-6: give:

replace by: With V=9 m3 this results in:

Page 10f3

page 4-7, 211d assumption, line 4: When the estimation for

a ...

will be a factor 3.16 to low.

replace by: When the estimation for

a

is a factor 10 too high, the ca1culated filter area will be

a factor 3.16 too high.

supplement: page 4-24, equation 4-55: The heat of crystallisation has been estimated

at 200 kJ/kg.

page 4-26, assumptions supplement and errata:

• The solubi1ity data taken from ChemCad are reliable.

• The mass fraction of solids in the outlet of the CDC can be up to 40 wt%. • A:H = 200 kJ/kg.

page 4-28, chapter 4.8.3, line 1: To keep the tube side ... chosen

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Errata and Supplements - FVO 3191

page 7-4, table 31: the datafor CaO are missing. supplement: CaO, 100 $/ton, 7165.90 ton/a, 0.72 M$/a.

Page 2 of3

supplement: The costs for the waste water treatment has been estimated. This estimation has

been based on information from "Hoogheemraadschap Delfland". The costs can be ca1culated with:

K

=

m .COD

ww 136 COD

=

Chemical Oxygen Demand (mg/l)

page 8-1, paragraph 4: POT 1.12 year and Ral 32.91 %

replace by: POT 1.12 year and ROl 33.91 %

Appendix VIII-3: figure i is missing.

Relation between Ss and Areq

18,20

~

,

~

18,00 17,80 17,60 N 17,40 .§.

N

17,20 17,00 16,80 16,60 16,40 0,5 0,6 0,7 0,8 Ss [-I

Figure i: Relation between Ss and Areq

---.

r---0,9

We discovered a major error in our extraction column design. These errata include a total

new version of page 4-23. A new calculation for the column dimensions can be found on this

(4)

t

FVO 3191 _{Equipment Design and Calculations} The column height, the column diameter and the falling velocity relative to the

column can be caIculated, with the flows through the column known, by sol ving the following equations: h = V drop,wall ' N . t_O.₇

D=~4Vtot

1t. h and v,EA E

=

---'"--- v,EA

+

v, water <l>VEA

V -v -v -v - '

drop,wall - drop EA - drop 0.25E1tD2 where: h height of column [m]

N number of real stages [-]

tO.7 time to reach 70% equilibrium [sJ

D diameter of column [m]

Ei fraction of phase i in total phase

Vdrop,wall velocity of drop relative to column [mis] Vlot volume holdup in column [m3], given by:

VEA velocity of ethyl acetate relative to column [mis]

v volumetrie flow rate [m3/s]

Looking at the volume flow rate through the column, consider that in the first two stages (at the top) the flow changes a lot by the transport of most of the TMP at those two stages. At stage 3 to 13 the volumetrie flow rate is nearly constant. Therefore the top stages need a larger diameter than the other stages. These diameters are caIculated apart from the other stages. The diameters and heigths are given in table 25.

Table 25:Dimensions ofthe column.

stage diameter (m) height (m)

1 0.42 4.54

2 0.32 4.54

3to13 0.29 49.99

The total height of the column can be calculated by:

Hi is the height of stage i [m]

For liquid removal and liquid injection, 2 m auxiliary column height is necessary at both column ends.

The tot al height of the column is estimated at 63.1 m.

4 - 23 [4-50] [4-51] [4-52] [4-53] [4-54]

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FVO Nr.

1

3191

1

"Fabrieksvoorontwerp"

Department of Chemical Process Technology

Subject

Production of20 ktonJa IMP (trimethylolpropane) from

n-butyraldehyde and formaldehyde

.

Autors

C.

(Kees) van de Beek

1. (Jurg)

Bremmer

Phone number

1.

W. (Jillis) Raadscheiders

M

.

(Martin) Reuvers

0172-215895

015-2621135

023-5371136

0180-411055

Keywords

IMP, 2-ethyl-2-hydroxymethyl-propane-l

,3-diol,

Polyhydric alcohols, Polyurethanes, Solvent extraction

Start Project :

Date Report :

11 October 1996

(6)

FVO 3191 Summary

Summary

A predesign of a 20 kton/a TriMetylolPropane (T!vfP) plant has been made. ,

The raw materials for the production were n-butyraldehyde and formaldehyde~ calciumhydroxide is used as catalyst and raw material. Ethyl acetate is used as extractant and solvent for TrvIP. The reaction section consists of three reactors in seritf over which the reactants are distributed.

Next the calcium formates are removed with sulfuric acid and the formed gypsum is removed with a rotary filter. The formaldehyde and formic acid is then stripped trom the process stream, the formaldehyde is purified and recycled. The TrvIP is then extracted trom the aquous phase with ethyl acetate and the T!vfP is crystalised with three, parallel, Cooling Disc Crystalisers (CDCs).

The solid TMP is removed from the ethyl acetate and the ethyl acetate is purified and recycled.

w

The solid TMP is >98 Yo pure. Production is 20.9 kton/year.

Byproducts are gypsum and formic acid, the latter needs further purification.

The economieal evaluation ofthe process resulted in a POT of 1.12 year and a ROl of32.91 %.

The economie feasibility for this proeess is very promising but the priee of T!vfP is of very high importance. As this plant would increase world production of T!vfP by 25 %, the pri~e of T!vfP is most likely to decrease. From the economie sensitivity analysis we learned that this has a large effect on the cash-flow.

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FV03191

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(8)

FV03191 7. Economy 7. 1. General Aspects 7.2. Investment Costs 7.2.1. Zevnik-Buchanan-Jansen 7.2.2. Taylor 7.3. Wages

7.4. Variabie Production Volume Costs 7.5. Total Costs

7.6. Depreciation 7.7. Interest 7.8. Gross Income

7.9. Calculation ofProfit and Cash-Flow 7. 10. Economical Criteria

7.10.1. Pay Out Time

7.10.2. Return On Investment 7. 11. S ensitivity Analysis

8. Conclusions & Recommendations 9. List of Symbols 9. 1. Abbreviations 9.2. Greek Symbols 9.3. General Symbols 9.4. Indices 10. Literature

Appendices

I. Process Flowsheet Il. Mass and Heat balance

lIl. Flowsheet Stream Compositions

-IV. Equipment Specifications V. Reaction Mechanism VI. The UNIF AC Model VII. "Chemiekaarten"

VIII. Information on Rotary Filter Design

IX Calculated Stage Compositions off all stages in the Extraction Tower X Economical Evaluation XI. Process Safety Contents 7 - 1 7 - 1 7 - 1 7 - 1 7-3 7-3 7-4 7-4

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-7-4 7-5 7-5 7-5 7 - 5 7-5 7-6 7-6 8 - 1 L \ 9 - 1 9 - 1 9 - 1 9-2 9-4 10- 1

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111

(9)

FV03191 Introduction

1. Introduction

The design ofthis TMP-plant is based on literature and patents. These dated back to even before the seventies. The renewed interest for TMP is boosted by the use of TMP in polymers, paints and high quality ink.

All articles and patents were closely examined and judged on their feasibility. We combined the most promissing parts of several patents and articles into a "new" process for the production of TMP. The separate parts are not new but the combination and the order in which they are applied are entirely new, as far as we know.

TMP can be used as a cross-linker in polymers. These cross-linkers are used in various products. Among these are paints, plastics, glues and, its most important use,

polyurethanes.Polyurethanes are formed by the polymerisation of di-isocyanates and polyhydric alcohols (figure 1).

TÎI?

nO = C = N -Rl - N = C = 0 + n

~

-+

Ic -

NB -Rl - NB - C - 0 -R2 -

ol

Lil

11

J

~

o

0

n

Fig. 1: Formation ofpolyurethane from a di-isocyanate and a di-hydroxy alcohol

The world-wide production ofTMP is now about 80 kton/~, the renewed interest jusitifies the design ofthis 20 kton/a plant which would enlarge the world production with approximately 25%.

The raw materials used for the production of TMP are formaldehyde and n-butyraldehyde.

Formaldehyde is a base chemical and available in abundant amounts. lt is available in a standard 37 wt% aquous solution stabilised with methanol. n-Butyraldehyde, which is produced from propene and carbonmonooxide,is also available in sufficient amounts.

The synthesis of TMP is relatively simple. The main problem in this process is the purification ofTMP from the reaction mixture. The purification has three main problems which will return in every purification step.

The first problem is that TMP is a solid at ambient temperature and pressure. lts melting point is 58° C. Just above this temperature TMP is a very viscous liquid. This makes transport of TMP through pipelines very difficult. Diluting TMP is a solution for this problem. TMP is soluble in water and various organic compounds. One ofthe raw materials for the synthesis is a aquous formaldehyde stream which allows the entire reaction to be carried out in an aquous solution.

The high solublity of TMP in both water and organic compounds brings on the second problem. In the purification section a liquid-liquid extraction is necessary. This requires,

because ofthe small differences in solubility, large amounts of extractant. To handle this problem, optimisation ofthe extraction conditions is necessary.

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FV03191 Introduction

The third problem is the formation of formic salts during synthesis. These compounds cause decomposition of Tl\1P at higher temperatures so these salts must be removed before any destillation or evaporation step can be performed. A solution for this problem is to remove the formates by ad ding sulfuric acid. Gypsum and formic acid are formed (see reaction 8). The gypsum can easily be removed with a filter, the formic acid by destillation.

To avoid chain-ending, most polymer ingredients/monomers have to be ultra-pure. However Tl\1P is tri-functional and most of its impurities, Dl\1P and di-Tl\1P, are bi- or four-functional so no chain-ending will occur. This allows a purity demand of ± 98%.

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The major producers ofTMP are Cilanese corporation Mexico (Hoechst) and Perstorp corporation Sweden

wW'

The economic feasibility of TMP jg proven during the last decades. TMP came into production in the early seventies and it still is. This plant is just an enlargement of the world production. This was made necesary because of new developments in the uses for TMP .

...-::--Though the process conditions are not extreme, the risks are not negl~ible. Formaldehyde and ethyl acetate are toxic and flammabie, so equipment has to be sealed to avoid leakage to the environment. Storage of37 wt% aq formaldehyde is problematic because ofthe formation of (polymeric) paraformaldehyde. Minimum storage temperature is 40°C. TIT (Just In Time) delivery can bypass this problem. Stabilisation with methanol is possible but not preferabie for the process. TMP itself ean eause dust explosions.

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FV03191 Starting Points

2. Starting Points

Starting point is the design of a plant for a TMP production of at least 20 kton a year with a purity of>98 wt%. Operation time is 8150 hours a year. The physical plant life is 20-25 years, because there are no high temperatures and pressures in the process. The economical plant life has been estimated at 15 years.

2.1 Basic concepts

A choice must be made between a batch- and a

continu~cess.

Because of the large production volume (»5 kton/a) a continue process has been chosen. A simplified pro ce ss flowsheet is given in figure 2.

ethyl acetate

I

n -butyraldehyde

~

ethyl acetate + TMP

C

TMP

-cJ

formaldehyde form

R

S

aldehyde calc iumhydroxide fo nnic acid sulfuric acid

wate r + heavy ends

gyps urn

Fig. 2: Simplified processflow sheet.

The raw materials and sulfuric acid are fed to a reaction section. The gypsum is separated from the reaction mixture with a rotary filter drum. In the separation section the remaining components are separated with the aid of several columns. One of these is an extraction column. The TMP is extracted from the aquous mixture with ethyl acetate. Finally the TMP is crystallised from ethyl acetate. The ethyl acetate is then recycled to the extraction column.

2.2 Reactions

In the following paragraphs the formation ofthe product and byproducts will be described. The reaction kinetics and enthalpys are discussed in chapter 3.

2.2.1. Formation of TMP from n-butyraldehyde and formaldehyde

The overall reaction ofthe synthesis ofTMP from n-butyraldehyde and formaldehyde is:

L

CH3 - CH2 - CH2 -CH~J 3 H2CO + Y2 Ca(OH)2 ~ CH3 -CH2 -C(CH20H)3 + Y2 Ca(HCOOh [1]

(12)

FVO 3191 Starting Points

This synthesis consists oftwo aldol condensations and one crossed-cannizzaro reaction.

First, one methylolgroup is added by the first aldol condensation and HAL (2-hydroxymethyl-l-butanal) is formed.

CH3 -CH2 - CH - CHO

I CH20H

(HAL)

The second methylolgroup is added by the second aldol condensation and DAL (2,2-dihydroxymethyl-l-butanal) is formed. CH3-CH2-CH-CHO + CH20 ~

I

CH20H CH20H I CH3 - CH2 - C - CHO I CH20H (DAL) [2] [3]

Finally, the aldehyde group is reduced by a crossed-cannizzaro reaction with formaldehyde to give the product TMP:

CH20H I CH20H I [4] CH3 - CH2 -C - CHO I CH20H

+ CH20 + ~ Ca(OH)2 ~ CH3 - CH₂- C - CH20H (IMP) + ~ Ca(HCOOh

I CH20H

The reaction mechanisms for reactions 2 to 4 are given in appendix V

2.2.2. formation of byproducts.

When the cannizzaro reaction takes place before both aldol condensations have completed DMP (dimethylolpropane) will be formed by the following reaction:

CH3 -CH2 -CH - CHO I + _CH20 + ~ Ca(OH)2 ~ CH3 - CH2 - CH - CH2 OH I (DMP) + ~ Ca(HCOO)2 CH20H CH20H _[5]

G~

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2 -2

(13)

Small amounts of other byproducts are formed by the reaction of two n-butyraldehyde molecules. These reactions give two new aldehydes: Xl (2-ethyl-3-hydroxy-hexanal) and X2

(2-ethyl-2-hexenal). C2Hs I 2 CH3 - CH2 - CH2 - CHO ~ 0.6 CH3 -CH2 -CH2 - CH - CH -CHO I OH C2Hs I 0.4 CH3 -CH2 - CH2 -CH = C - CHO (X2)

Xl decomposes into two n-butyraldehyde molecules by reaction [7]: C2Hs

I

CH3 -CH2 - CH2 -CH - CH -CHO ~ 2 CH3 - CH2 -CH2 - CHO

I

OH

A crossed-cannizzaro reaction between two formaldehyde molecules is also possible:

Finally, dimerisation ofTMP gives another byproduct, di-TMP:

CH20H I 2 CH3 - CH2 - C - CH20H I CH20H ~ CH3 -CH2 - C - CH2 - 0 - CH2 - C - CH2 - CH3 I I CH20H CH20H

2.2.3. Conversion of calciumformate to formic acid

(di-TMP)

[7]

[8]

[9]

The ealciumformate formed in the erossed-eannizzaro reactions [4], [5] and [8] is converted into calciumsulfate (gypsum) and formic acid by the following reaetion:

[10] [6]

(14)

2.3. Specification of raw materials and produets

The specifications of raw materials, products and utilities are given in tables 1 to 3.

Table 1: Specifications ofraw materials and produets

Component purity impurities

n-butyraldehyde1 99 wt% water (0.5 wt%)

acidity 0.3 wt%

remarks

1988 world production about 4.4 Mton/a

formaldehyde 37 %aq 0.02 wt% acidity 37 wt% aquous solution, stabilised

with 1 wt% methanol Ca(OH)2

sulfuric acid

EtAcl

formic acid

50%aq solution in water 50 wt%

gypsum TMP 98 %mol 99wt% 85;90; 95; 98 or >99 wt% >99wt% 98wt% water water water water DMP(0.7 wt%) other org.0.2 wt% EtAc (1.1 wt%)

1 For the simulations 100 wt% pure n-butyraldehyde and ethyl acetate have been used. The identity ofthe

impurities were unknown at the time ofthe simulations.

Table 2: Specification of utilities

Utility Specifications

Low pressure steam P=3 bar, T= 190°C, T cond= 13 3.5 °C

Electricity 220 V, 380 V

Cooling water T in=20 °C, T out= 40°C

Heavy Ends, Byproducts Waste treatment

Table 3: Physical properties.

Component Molweight Bp Mp density

[gImol] [0C] [0C] [kgldm3 @T] 74.8 -96A 0.802

®

-15 1.1 580 2.24 n-butyraldehyde 72.12 formaldehyde 30.03 Ca(OHh 74.09 pnce [$/kg] 0.947 OA18 EtAc 88.1 77 formic acid 46.026 100.56 methanol 32.042 64.7 CaS04 136.14 -83 8A -97.68 0.9 1.2 @20 0.8 @20 1.256 0.903

{~\~~

TMP 134.18 289 58 1.08-1.1 2.62

---.J

Because ofthe problematic formaldehyde storage a plant location near a formaldehyde plant is

strongly recommended.

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(15)

FVO 3191 Process Structure and Process Flowsheet

3. Process Structure and Process Flowsheet

3.1. Overall

process

description

Formaldehyde, n-butyraldehyde and calcium hydroxide are fed to the reaction section. The temperature in the reactors is 58°C and the pressure 3 bar. The entire formaldehyde feed (7) is added to the first reactor (R5), the other raw materials are divided between the three reactors (R5, R6, R7). In this way, a formaldehyde excess is maintained, which results in a higher conversion and selectivity. This design has been based on lit. 21.

The distribution ratios of the n-butyraldehyde and calcium hydroxide, as weU as the reactor temperature, were optirnised with ChemCad simulations. As starting ratios for the

optimisation data from a patent (Iit 21) were used.

In reactor R8 the calcium formate is converted to calcium sulfate by ad ding sulfuric acid (20). The calcium sulfate is removed with a rotary filter (M9). In tower TlO the light ends

(formaldehyde and forrnic acid) are separated from the process stream. In tower TIl, the formaldehyde is purified and recycled to the reaction section (43). Formic acid can be purified to 85 wt% with a cascade ofthree destilation towers. (lit 22 & 23)

The process stream (28) is fed to the extraction column (TI2). The TMP is extracted from the aquous stream with ethyl acetate (52). The column temperature is 85°C, the column top pressure is 3 bar. The ethyl acetate stream from the extraction column (48) is cooled down to 65°C with a recycle steam ethyl acetate (59) from the band filter (M14).

Subsequently, the TIMP is crystallised from the ethyl acetate in three paralel CDCs

(M13A,B,C; Cooling Disc Crystalliser). In the crystallisers the ethyl acetate is cooled down to 30°C. At this temperature, the ethyl acetate is just saturated with byproducts, so a good purity / recovery ratio is achieved. The solid TMP (67) is separated from the ethyl acetate (57) with a band filter (M14). The ethyl acetate is purified in tower T15 and recycled (62). The heavy ends are purged to the waste treatment (66).

Many possibilities for the purification section of the plant were found in the literature. Most were rejected during discussion but the most prornising ones were, excluding those used, an extraction with xylene and n-butanol [lit. 38];. this was rejected because oftechnical

difficulties and the environmental etfects ofxylenes. An evaporation ofTMP was rejected because this was done with the formic salts still in the reaction mixture. More recent patents discussed the decomposition of TMP at higher temperatures caused by the presence of formic salts.

An ion exchange column for the removal ofthe formates was rejected because ofthe high costs [lito 34].

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(16)

FVO 3191 Process Structure and Process Flowsheet

3.2. Reaction kinetics

A literature search for kinetic data resulted in (only) two articles, which were, above all, written in Czech. The kinetic data are given in table 4.

Table 4: Kinetic data and reaction enthale.Y

reaction kO Ea LlRH k at 58°C (=331.15 K)

number ~kJ/mol) k=kO · eXE[-EAfRT]

1 -137.21 2 6.35332* 1012 7.663429* 1 04 -24.81 5.182 3 2.83364*1011 7.02782*104 -17.26 2.325 4 1.00808* 1017 10.385*104 -95.13 4.187 5 1.54711 * 1013 8.56076* 1 04 -275.62 0.485 6 1.026* 1013 7.4826*104 -104.33 16.139 7 2.457* 1012 7.4826*104 0 3.865 8 4.6038*105 6.2229*105 -96.29 7.03e-5 9 0

The dimerisation of TMP to di-TMP is not mentioned in this kinetic model. Because it is mentioned in several artic1es and patents we did take account of it. We assumed a 2% overall conversion of TMP to di-TMP.

The conversion ofn-butyraldehyde to HAL and DAL (aldol condensation) is catalysed by OH-. This effect can not be simulated in ChemCad.

3.3. Thermodynamics

Most organic compounds in the model mentioned above had to be introduced in ChemCad.

Xl, X2, HAL, DAL, DMP, di-TMP were introduced with the group contribution method of

the UNIF AC model. Therefor this thermodynamic model was used for the entire simulation.

Because ofthe difficulties simulating liquid-liquid extraction with ChemCad the extraction column (T12) was calculated with MS Excel. The units calculated with MS Excel were also based on the UNIF AC model. The crystalliser (MB) and the rotary filter (M9) were also calculated with MS Excel.

Several T,x,y diagrams ofthe binairy systems containing an azeotrope, are given in appendix VI. The enthalpy calculations were based on the information provided by the model in lit. 12

,extra components were added.

(17)

FV03191 Process Structure and Process Flowsheet

Enthalpy model: H=A+B-T

Table 5: model coefficients for enthalpy calculations

Compound A [kJlkmol] B [kJ/kmo!. 0c]

*106 *102 n-butyraldehyde -0.25684 l.58398 formaldehyde -0.17784 0.91692 HAL -0.45949 2.50090 DAL -0.65459 3.41783 calciumhydroxide -l.01180 0.84638 TMP -0.71291 3.26180 calciumformate -0.14408 2.92880 methanol -0.23758 0.80751 DMP -0.69848 2.37664 water -0.26319 0.75266 XI -0.51368 3.16796 X2 -0.51368 3.16796 formic acid -0.378600 1.00000

(18)

FV03191 Equipment Design and Calculations

4. Equipment Design and Calculations

4.1. Pump design

This paragraph describes the design of unit P2, increasing the n-butyraldehyde feed stream pressure from 1 to 3 bar. Pump selection is based on the flow rate, required he ad (hman) and

the physical properties ofthe fluid being pumped. The small flow rate, required head and the low viscosity of the fluid justify the use of a centrifugal pump.

The isentropic power needed has been calculated with:

LW

pp•isen

=

-Wp.isen

=

<Pm'

p

=

<Pv' LiP [4-1]

The real pump power (Pp) depends on the pump efficiency. The pump has a mechanical and

hydraulic efficiency. [4-2] where: llm llh IIp mechanical efficiency [-] hydraulic efficiency [-] overall pump efficiency [-]

The mechanical and hydraulic efficiency are not known, but they are replaced by an overall pump efficiency. This efficiency depends on the flowrate and the pump used. An overall efficiency of 40% has been estimated from fig. 10.62 [lit. 3]

A check on the NPSH-available (net positive suction head) must be made. It is important that the pump suction pressure does not fall below the vapour pressure of the liquid being pumped, otherwise cavitation may occur. This can lead to a reduction in the flow or cavitation in the pump. The available suction head is determined by the suction piping design and the vapour pressure of the fluid. If the effect of the piping design is neglected a simple relation between NPSH and vapour pressure follows:

where: Pin

Pv g

suction pressure [Pa] vapour pressure [Pa]

gravitational acceleration = 9.81 [m1s2]

[4-3]

(19)

The results of equation 4-1 to 4-3 are given in table 6

Table 6: results a/pump calculatian

Pump P 2 value units INPUT <l>v 2.2697 m3/h p 800 kg/m3 M 200 kPa Pin 100 kPa Pv 11.7 kPa llr 40 % OUTPUT Pp,isen 126.094 W Pp 315.236 W NPSHavail 11.2513 m hman 25.4842 m

Since n-butyraldehyde is not corrosive, simple construction materials, such as stainless steel, cao be used.

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4.3. Reactor Design

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er (i.e. the diameter equals the height),

ith the current flow, tbis leads to a

The equations 4-4 to 4-6 define the reactor dimensions.

~

.v D = 3 _ r n [4-4] [4-5] A =ns ·D·H [4-6] glVe: Dr 2.25 m diameter reactor H 2.25 m height reactor Ag 4 m2 (ground area)

(20)

The total heat ofreaction in this reactor is calculated in a energy balance and is 333 kW. The cooling ofthe reactor can be done by ajacket or intemal coil cooling. To determine whether the shell cooling is possible equation 4-7 is used to ca1culate the required shell surface. The heat transfer coefficient U is estimated at 400 W/m2 K

An effective cooling area of 90 percent of the outer shell was assumed. The total cooling surf ace area required can be ca1culated with equation 4-7:

Q

Areq

=

U·ÓT This results in A = 29.73 m2

[4-7]

The surface area ofthe shell is too small for jacket cooling, therefore cooling with an intemal coil is used. When an intemal coil with a outer diameter of 5.08 cm (=2 inch) is used, the dimensions of the (cu bic cylindric) tank and coil are given in table 7

Table 7: Tank and cai! dimensians

Tank dimensions Coil dimensions

D 2.29m D 5.08 cm

H 2.29m L 186.3 m

Veff 9.00 m3 V 0.38 m3

Total tank volume is 9.38 m3

Agitatian

The agitation is done with a turbine agitator. The following standard dimensions were chosen. DalDt =0.33 plDa = 1 where Da Dt p agitator diameter [m] reactor diameter [m] blade pitch

With the ca1culated reactor diameter these equations give Da = 0.77 and p = 0.77

The Reynolds number is calculated by equation [4-8] p·N·D2

Re= a

11

where N = 1 S-I (agitator speed, revolutions per second)

Re = 1. 13.106

[4-8]

(21)

-

---~~~~~~~~~~~-FV03191 Equipment Design and Calculations

From fig. 10.58 in Coulson and Richardson [lit. 3] we find the power number Np.

Np = 0.33

With equation 4-9 the shaft power can be calculated. N

=

Pshaft

p D~N3p

[4-9]

This results in P shaft = 71.9 W

4.4. Design of Rotary Filter

T 0 separate the gypsum from the reaction mixture, we first designed a hydrocyclone. After a

conversation with Prof. B. Scarlett M.Sc. the designed separation proved irrealistic. In our design we assumed an underflow with 70 wt% solid, taken from Bradley [lit.15]. This book dates back to 1965, and our assumption seemed to be incorrect.

According to Mr Scarlett a underflow solid fraction of 10, maybe 20 wt% is arealistic value. Indeed, hydroclyclones are used more often as a classifier than as a particle separator. We had to find an other solution for this separation step.

From Mr Witkamp, an expert on gypsum separation, we leamed that the PSD (particle size distribution) would likely be an exponential distribution with a number average particle size of somewhere between 50 and 100 J...I.m, 60 J...I.m was assumed as number average size. Of course this is dependent on rate of formation of the gypsum crystals and the crystal shape. Such relatively large particles can easily be removed with a rotary filter.

A rotary filter normally filters a suspension from a trough. A pressure drop of 0.5 bar is sufficient. Because our process fluid has a pressure of 2.7 bar, a cover is placed on the filter drum to keep the pressure on the outside ofthe drum at 2.7 bar and the inside ofthe drum at 2.2 bar.

,

-

- '"

._, cover

(22)

FVO 3191

The design of the rotary filter drum starts with the Ruth equation:

where: Il a f <1>"

liquid viscosity [Pa.s]

specific cake resistance [m2/kg]

a function of the slurry composition [kg/m3 tot a! volume of filtrate per unit area [m3 /m2]

filter medium resistance [-]

Equipment Design and Calculations

[4-10]

R

t time of cake on filter being submerged in suspension [s] This equation is derived from Darcy 's law (see appendix VIII).

The fin equation [4-10] is ca1culated by:

where PI Ss Sc f= PI I-S _s 1-S _c liquid density [kg/m3]

solid fraction in slurry [kg/kg] solid fraction in cake [kg/kg]

[4-11 ]

The rate of formation of the cake depends on the residence time of an element of the drum in the suspension. Consider an element dA of the drum. The residence time can be calculated as follows: where cp N cp 1 t = _ · -2n N

angle of subrnerging [rad] rotations per second [S-I]

[4-12]

Equations 4-11 and 4-12 can now be substituted in equation 4-10 to calculate <1>". But each element dA participates N times per unit of time, therefore the total throughput equals:

<1>v = N·A<1>" [4-13]

In this design the filter medium resistance is neglected (R « a) with regard to the specific cake resistance. Equation 4-10 becomes:

[ 4-14]

(23)

The specific cake resistance was estimated at 1011 m2fkg and the diameter ofthe filter at 2 m. The filter rotates at 1 rpm and the angle of submerging is 120°. This means that the di stance between the horizontal axis ofthe drum and the surface ofthe slurry is 0.58 m.

The relations between <p, D and h are given in appendix VIII

The solid fraction in the slurry is calculated from the feed stream composition. The solid fraction in the cake is estimated at 0.98.

The total filter area required can be calculated byequation [4-13].

A = 16.36 m2

When the diameter of the filter is 2 m the length can be calculated. The results of the calculations are given in table 8.

Table 8: Resu/ts ofrotaryfilter calculation

value dimensions ~'t.J.OO..:n<····

..

Yj a. M> cp Ss Sc PI N <1>v D 2.00 m ···Q'tWt'~lJj; }···U·.··.·<···:··:·::::::::::···:·:·:·

...

.

f 178 m2/kg t <1>" A L 20 0.0128 16.36 2.60

(24)

Assumptions.

• The filter medium resistance is neglected with regard to the specific cake resistance.

• A specific cake resistance of 1011 m2/kg was estimated. The calculated filter area is

dependent on the specific cake resistance. Equations 13 and 14 can be rearranged to 4-15

A 2

=

<1>~llaf

req 2N2 LiP. t [4-15]

The filter area required is proportional to the square root of a. When the estimation for a is a factor 10 too high. The calculated filter area win be a factor 3.16 to low. So a good estimation for a is essential. Because we first designed a hydrocyclone we did not have enough time for a literature search for a good estimation. Therefore we took a value for a mentioned in an excercise in "Deeltjestechnologie" [lit. 17]. The particles in this excercise were relatively large, like the gypsum particles in this design.

• The solid fraction in the cake is estimated at 0.98. This means that the solid product (gypsum) contains only 2 % liquid. The calculated filter area is dependent on the liquid fraction in the gypsum product. From figure i in Appendix VIII, we learn that a change in solid fraction will hardly cause a difference in filter area required. In fact the solid fraction in the cake is less than 0.98, but after washing and drying a value of 0.98 can be reached.

• For further calculation no liquid in the gypsum is assumed.

4.5. Destillation Column Design

The design ofthe formaldehyde formic acid separation column (TI2) is dealed with in this chapter. The column is designed, following the methods described in Coulson & Richardson' s Chemical Engineering, Volume 6 [lit. 3]. When no source is specified for a page number or a figure it is taken from lit. 3.

To design a column, a heavy key and a li h e components used for this column are ethanol heavy key component.

component have to be choosen. The key s light key component and formic acid as juC'

s6

(25)

In the first place the minimum number of stages (at total reflux) is estimated with the Fenske

equation:

I

[XLK][XHK ]

N=Og~d~b

m 10g(a

LK)

[4-16]

where:

XLK

mol fraction ofthe light key component

XHK mol fraction of the heavy key component

Nm minimum number of stages at total reflux, including reboiler

aLK relative volatility of the light key with respect to the reference component.

d refers to distillate

b refers to bottom product

To estimate Nm, specifications ofthe top and bottom molefractions ofthe key components are

needed. The specifications used are given in table 9.

Table 9: Specifications Component Methanol Formic Acid Xi,d 0.050 0.031 Xi,b 0.002 0.041

The relative volatility is calculated by equation 12.37, Smith and van Ness [lito 19]:

y.psat

a ..

=

1 1 IJ Y psat

J J

activity coefficient of component i

Vapour pressure of component i

F or all aij the heavy key is used as component j.

Due to lack of experimental data the activity coefficents had to be calculated by the

(predictive) UNIFAC model. The equations for tbis model are given in appendix VI.

The vapour pressure is calculated by:

temperature [K]

[4-17]

[4-18]

where T

A,B,C,D,E component dependent parameters (extracted from the CHEMCAD

(26)

FV03l9l Equipment Design and Calculations

The top and bottom pressure of the column are estimated. With these pressures, the

temperatures of the feed, top and bottom are calculated. A bubble point calculation was made to determine the top temperature and dew point calculations we re done to calculate the feed and bottom temperatures (Smith and van Ness, p 381-393, lit. 19). For these calculations an estimation was made (from CHEMCAD results) ofthe total composition ofthe top and bottom streams. The estimations and results are given in table 10.

Table i 0: Estimations of compositions and pressure, and calculated temperatures

Feed Top Bottom

T (OC) 60 117 129 P (bar) 2.50 2. 46 2.60 N-Butyraldehyde 0.0002 Water 0.90 Formaldehyde 0.047 Methanol 0.009 Formic Acid HE 0.039 0.0002 0.0012 0.63 0.28 0.05 0.031 0.00005 0.00001 0.96 0.00009 0.0015 0.041 0.0002

HE are the heavy ends. These are mainly Xl, X 2, DAL and HAL. During this design the

properties of these components are assumed to be the same as the properties of TMP (due to lack of (experimental) data). The fractions given are mole fractions.

With these calculated temperatures, the vapour pressures and from them, the relative

volatilities can be calculated (eq. 4-17). The relative volatilities of the light key component at the top and at the bottom of the column are given in table 11

Tableii: Results of calculations

component Pd [Pa] Pb [Pa] Xd Xb aij,d aij,b Methanol 583244 820973 0.050 0.002 9.9 6.4 Formic Acid 159858 221898 0.031 0.041 1.0 l.0 Finally Nm can be estimated with the F enske Equation. (eq. 4-16)

Nm = 1.8 stages

When Nm is known, the minimum retlux ratio is calculated by the method of Colbum and

Underwood:

[ 4-19]

where:

R.n

e

the minimum reflux ratio the root of equation 4-20:

(27)

FV03191 Equipment Design and Calculations where: f q ~ a i)-xi f L.J ----,---,',--

=

1 - q a _IJ-8 refers to feed

depends on the condition ofthe feed:

Heat of Vaporisation of the feed q = Heat contained in the feed

[4-20]

[4-21]

If the feed enters the column as a liquid at its boiling point the value of q is 1. The value of 8 must lay between the values of the relative volatility of the light and heavy key components, and is found by trial and error. The resuIts ofthe calculations are given in table 12

Table 12: Resu/ts of calculation of Rm

Rrrt

0.0002

q 0.99

8 7.5

The number of ideal trays for the column is estimated by the method of Erbar and Maddox. The reflux ratio is choosen at 0.9. This resuIts in a number of6.9 ideal trays (read from figure 11.11).

The feedpoint location is estimated by using the emperical equation given by Kirkbride:

[4-22]

where: Nr number of stages above the feed, including any (partial) condensor Ns number of stages below the feed, including the reboiler

B molar flow bottom product (562 kmole/h) D molar flow top product (111 kmole/h) LK refers to light key

HK refers to heavy key

When the ratio NriNs and the total number of stages are known, the position of the feedstage can be calculated.

NrlNs = 0.532

Feed stage (ideal)

=

2.4

The stages are ideal stages. The real feed point location is deterrnined by deviding the calculated feedstage by the stage efficiency (when the efficiency is constant in the whole section).

(28)

The distrubution of the non-key components can be estimated with the method of

Hengstebeek and Geddes:

where A, C are parameters, fitted by the distribution of the key components

[4-23]

When D and B ofthe key components are known, A and C can be calculated. Then the distribution ofthe non-key components can be calculated using A en C in the Hengstebeek and Geddes equation. The results of this calculations are given in table 13.

A = -0.8257 C = 1.8408

Table 13: Distribution ofnon-key components

Component Feed Di Bi xi)" Xi,d Xi,b

[kmolJh] [kmolJh] [kmolJh] [-] [-] [-]

butyraldehyde 0.14 0.13 0.01 0.000 0.001 0.000 water 608.93 72.86 536.07 0.904 0.646 0.956 formaldehyde 3l.63 30.90 0.73 0.047 0.274 0.001 methanol 6.36 5.55 0.82 0.009 0.049 0.002 forrnic acid 26.28 3.42 22.87 0.039 0.030 0.040 HE 0.10 0.00 0.10 0.000 0.000 0.000 total 673.44 112.85 560.59 1.000 1.000 1.000

These distributions are approximately the same as the estimations. The estimations have to be revised until the difference between the calculations and the estimations is acceptabely low.

4.5.1. Approximate column sizing

For the column the molar flows ofthe vapour and the liquid in each section ofthe column are assumed to be constant. The only significant change in molar flowrate will be at the feed stage. Here the liquid flow changes while the vapour flow remains constant. So the column has to be designed in two sections.

Plate spacing

F or the further design of the column the plate spacing has to be choosen. Here a plate spacing oflt=0.45 mis used. The approximate column sizing is done by the method described in

Coulson & Richardson page 509

(29)

FVO 3191

Column diameter

The column diameter can be calculated with:

where: De Vw

Pv

Uv

column diameter (m)

maximum vapour rate (kg/s)

vapour density (kg/m3) (calculated with ideal gas law) maximum allowable vapour velo city (mis)

The maximum allowable vapour velo city is calculated with:

where: lt PL 2 PL - Pv [ ( ) ] 0.5 Uv

= (-

O.l71·lt +0.27·lt -0.047)· Pv plate spacing (m) liquid density (kg/m3)

Equiprnent Design and Calculations

[4-24]

[ 4-25]

The densities and several other physical properties are ca1culated with equations and

parameters extracted from CHEMCAD. The average values for each section are given in table 14.

Table 14: Physical properties olthe column streams

Top Bottom section section PL (kg/m3) 934.3 954.8 pv (kg/m3) l.59 1.42 M (kg/kmol) 2l.96 19.51 Vw (kg/s) l.36 l.13

Lw

(kg/s) 0.62 4.20 aL ili/m) 0.05 0.05

Using the values given in table 14 the values ofuv and De can be calculated. This results in the

values listed in table 15.

Table 15:Estimation ol diameter

Uv (mis) De (m) Top Bottom section 0.9653 l.0616 section l.0299 0.9928

(30)

FV03\9\ Equipment Design and Calcu\ations

4.5.2. Plate design

The column is equiped with sieve plates because of the low costs of this plate type, the good

performance for this kind of application and the low pressure drop. The plates are stacked

plates because ofthe small diameter (less than l.2 m)

Column diameter

The column diameter is again calculated. Here the diameter is based on flooding

considerations.

D

=

~4~c

Ac

is the column cross-sectional area calulated with:

Area required for flow (Areq)

A=....,---'----'-'---c (100 - percentage downcomer area)

As fi.rst trial a downcomer area of 12% of the total area is assumed.

<pffiax

A = _v_

req U

v

<t>V,max is the maximum volumetric vapour flow rate (m3/s) calculated from:

where: V M D R Uv <pffiax _ V·M v -3600· Pv

molar flowrate = D·(R + 1)= 213.7 kmoVh

average molar weight [kg/kmol] distilate rate [kmol/h]

reflux rate [-]

actual vapour velo city, 85% offlooding velo city (Uf) [mis]

The flooding velocity is calculated with:

Where:K

aL

u

f

=

Kj

.(~)"2

.

~PL

-Pv

0.02 Pv

a constant depending on the flow factor (FLV)

surface tension of the liquid stream [N/m2]

[4-26] [4-27] [ 4-28] [4-29] [4-30] 4 - \3

(31)

FVO 3191 Equipment Design and Calculations

The flow factor can be calculated with:

[4-31]

When FLV is known the value of KI can be read from figure 11.27. The calculated values are

given in table 16

Table 16: Calculation results

Top Bottom section section FLV [-] 0.019 0.14 Kl [-] 0.080 0.070 Uf [mis] 2.32 2.19 Uv [mis] 1.51 l.86 <1>vmax [m3/s] 0.82 0.81

Are

q [m 2 ] 0.54 0.44

Ac

[m2] 0.61 0.50 D [m] 0.88 0.79

Liquid flow patern

For the determination ofthe flow pattern the maximum volumetric liquid flow rate has to be calculated. This can be done by using equation 4-29, substituting each parameter by the relevant parameter for the liquid flow. (for example: V ~L (L= Lw·3600))

From the calculated values and figure 11.28 the flow pattern ofthe different sections can be determined:

Top section Bottom section

<DL max= 0.0007 m3/s

<DLmax= 0.0044 m3/s

Crossed flow (Single pass)

reverse flow pattern cross flow (single pass)

Reverse flow

• Flow pattern

(32)

FV03191 Equiprnent Design and Calculations

Provisional plate design

The plate design for the bottom section is summarized in table 17, for the top section the design summary is given in table 18.

Table i7: Plate design of bottom section

Column diameter De 0.79 m

Column Area A, 0.50 m2

Downcomer Area A.! 0.06 m2

Net Area Art 0.44 m2

Active Area

Aa

0.38 m2 Hole Area ~ 0.04 m2 Weir lenght lw 0.60 m Weir height hw 50.00 mm Hole diameter dh 5.00 mm Plate thickness tg 5.00 mm

Table i8:Plate design of top section

Column diameter De 0.88 m

Column Area A, 0.61 m2

Downcomer Area A.! 0.04 m2

Net Area _Art 0.58 m2

Active Area

_Aa

0.54 m2 Hole Area ~ 0.05 m2 Weir lenght lw 1.01 m Weir height hw 50.00 mm Hole diameter Dh 5.00 mm Plate thickness

_tg

5.00 mm

(12% of column area, chosen) (A,-Al)

(A,-2A.!

=

Art-Al)

(10% of active area, chosen) Figure: A.!/A,

=

12 ~ IwlDe

=

0.76

(6% of column area, chosen) (A,-A.!)

(A,-2A.!

=

Art-A.!)

(10% ofactive area, chosen)

lw is estimated by 0.5*lw(base)+0.8*De

This design has to be checked on weeping, pressure drop, downcorner back-up and entrainment. These are described below.

Weeping

To check on weeping, the minumum vapour velo city trough the holes has to be determined with:

minimum vapour velo city through the holes (based on hole area) (mis)

hole diameter (mm)

a constant dependent on the depth of clear liquid on the plate (hw+how)

[4-32]

(33)

The value ofthe weir crest (how) can be calulated from the Francis weir equation:

[ 4-33]

The value ofK2 can be read from figure 11.30.

The lower limit of the operation range occurs when liquid leakage through the holes becomes excessive. The vapour velo city at this so called weep point, is the minimum value for stabie operation. Here the lower limit ofthe operation range is assumed to be located at 70% ofthe maximum liquid rate. From this value the lower limit ofthe vapour velo city through the holes can be deterimined. The calculated values for this column design are given in table 19. The actual minimum vapour velo city through the holes is calculated by dividing 70% of the actual maximum vapour rate (uv) by the total area ofthe holes (An).

Table 19: Resu/ts ofthe weeping check.

Top section Bottom section

actual mIrnmum actual illIrumum

Lw

(kg/s) 0.43 2.94

how (mm) 4.45 22.21

hw+how (mm) 54.45 72.21

K2 30.15 30.15

Uh (mis) 11.77 9.35 15.07 10.24

From table 19 it follows that the actual minimum operating rate will be above the weeping point.

Plate pressure drop

The pressure drop per plate can be calculated from equation 4-34:

[ 4-34] Where ht (total pressure drop per plate (mm liquid)) is calculated from:

[4-35]

(34)

The results ofthe pressure drop calculatations are given in table 20. Table 20: Plate pressure drop

Top Bottom section section Uh (mis) 16.82 21.52 Co (-) 0.84 0.84 ht (rom liquid) 103.8 141.4

M>

~aî 951 1324 Downcomer design

The downcomer baclrup is given by:

where: hb

Lwd

Am

downcomer baclrup (rom liquid) liquid flowrate in downcomer (kg/s)

smalle st of

AI

and

Aap

(clearence area under downcomer) (m2)

The clearence area under the downcomer is given by:

For a safe proces design the following mIe has to be obeyed:

The residence time, given by equation 4-39, should be at least 3 seconds.

Where tr is the residence time. [s]

The results of the calculations are given in table 21. Table 21:Downcomer liquid baclcup results

top bottom section section

Aap

(m2) 0.04 0.02 hb (m) 0.16 0.23 0.5'khw (m 2 ) 0.03 025 tr ~s) 8.94 3.04 [4-36] [ 4-37] [4-38] [4-39] 4 - 17

(35)

Recheck entrainment

The entrainment can be estimated with the Fair correlation between the fractional entrainment

\jI (kg/kg gross liquid flow) and the Flowfactor FLv, with the percentage approach to flooding as a parameter. \jI can be determined from figure 1l.29. The percentage flooding is given by:

v~

percentage flooding = Uv = An _[4-40]

U_f U_f

\jI has to smaller than 0.1 for a safe design. The results ofthe entrainment check are given in

table 22.

Table 22:Results of entrainment calculations

top section bottom section

Per cent 0.61 0.85

flooding

Fiv (-) 0.0187 0.1425

\jI~-) 0.09 0.025

For both sections \jI is smaller than 0.1 so both sections are safely designed with respect to

entrainment.

Plate efficiency using Van WinkIe 's correlation

The plate efficiency is determined for both sections using Van WinkIe 's correlation:

r

L-\-\

t

~

[ J

O

.

08

0.14 0.25

0.07 cr L /J.L hw . Uv . Pv

C

L

uJ

(PL

D

J

~

L

(A

~

J

where: DLK liquid diffusity, light key component (m

2 /s) estimated: 9-10-9 m2/s Top section: Bottom section:

Emv

= 0.66

Emv

= 0.64 [4-41]

The real number of stages can now be calculated bij deviding the number of ideal stages by the efficiency.

Nrcal

= 11 stages

. top section

bottom section feed stage

4 stages 7 stages stage 5

(36)

FV03191 Equiprnent Design and Calculations

Now the pressure drop over the column can be calculated:

LV> column LV> top section LV> bottom section 0.14 bar 0.04 bar 0.10 bar

The calculated pressure drop is approximately the same as used in the estimations.

Column height

The height of the column can be estimated by:

Hcolumn

=

(N column - 1) ·1 t + 1.5 + 2.5 + 2 [4-42]

The values given in this equation are for the height above the top plate (needed for the vapour disengagement and mist elimination), the height below the bottom tray (gas inlet and liquid dis charge ) and a skirt. However this equation is for columns with a diameter above 1 m, this equation is used by absensce of better equations.

For this column the height wiU be Hcolumn = (11-1)*0.45+ 1.5+2.5+2 = 10.5 m.

Assumptions

• The molar vapour flow is constant over the whole column. • The molar liquid flow is constant in each column section.

• The physical properties of the heavy ends are the same as these of TMP. • The plate efficiency is constant in a whole section.

• Relative volatility of water is 0.95 throughout the whole column (due to obvious calculation results caused by azeotropes)

• Equation 4-42 can be used to estimate the column height.

4.6. Design of Extraction Column

Because ChemCad had some fata! difficulties with the simulation ofthis liquid-liquid

extraction, the extraction column T12 was calculated with MS Excel. The Kremser equation is

used to calculate equilibrium between water and ethyl acetate and the transfer of TMP. The mass fraction not removed TM.P is given by:

where: N

Tl

S

f

=

--::-::S_--:-:-I_ SCNoT]+l) -1

theoretical number of stages

[4-43]

stage efficiency (assumed to be 0.7, so 70% equilibrium is reached on a stage) Separation number, given by equation 4-44.

(37)

FV03191

k·V

s =

-L

mass flow ethyl acetate [kg/h] mass flow water [kg/h]

Equipment Design and Calculations

[4-44]

where: V

L

k equilibrium constant ethyl acetaat/water. k=2.8 (patent, lit 22)

To ca1culate the real number of stages the fraction TMP not removed was chosen at 1.0.10-9.

The real number of stages (= Nlrl) needed is about 13.

Now the compositions ofthe streams out ofthe tower can be estimated with stage to stage ca1culations. Therefore the following assumptions were made:

• On each tray 70% ofthe equilibrium is reached. This means that when an amount ofTMP

is to be transferred from the water to the ethyl acetate stream on one stage, according to the amounts of water and ethyl acetate and the equilibrium constant on that stage, only 70% ofthat amount is actually transfered.

• The equilibrium constants for TMP and the heavy ends are extracted from a patent [lito 22]. The other k-values were estimated by:

k

=

solubility of component in ethyl acetate

so lub ility of component in water

The equilibrium constant used for TMP: k = 2.8

The equilibrium constant used for the heavy components: k = 25

The equilibrium constant used for formic acid: k=0.59

The equilibrium constant used for sulfuric acid: k=0.05

There is no transport of other eomponents.

[ 4-45]

• The amount of ethyl acetate dissolved in water and the amount of water dissolved in ethyl acetate is, with an efficiency of 70% per stage, the maximum solubility, saturated. The solubilities were extracted from lit. 40.

• The neeessary mass flow of ethyl acetate was twiee the waterflow, aeeording to the patent

[lit. 36 & 22]

With all these assumptions stage to stage ca1culations have been done. This resulted in the

(38)

Table 23: Stream comeositions.

Stream top in top out bottom in bottom out

k~h k~ k~h k~h water 1622 188 1310 1434 ethyl acetate 0 3129 3244 115 TMP 2955 2955 0 0 heavy ends 330 330 0 0 formic acid 9 9 7 7 sulfuric acid 0.03 0.001 0.0004 0.03 waterphase 4577 1549

eth~l ac. Ehase 6272 3244

Due to recycling of ethyl acetate several components are already dissolved

in the bottom-in stream. The recycle of ethyl acetate reduces operating costs for tbis column

but does not effect the efficiency of the column.

To calculate the dimensions ofthe column the height and diameter have to be estimated. The height ofthe column can be calculated from the time needed to re ach 70% of equilibrium state (because of70% efficiency) and the velo city ofthe falling water droplets in the ethyl acetate.

~

Water droplets in ethyl acetate have been chosen because ofthe greater amount of ethyl

acetate per amount water and because tbis situation gives the best contact between TMP

dissolved in water and the extractant.

The diameter of the drop lets is taken at about 3 mmo [lit. 41].

Calculation of time to reach 70% equilibrium.

Mass transport can be calculated with:

where: V A c

EA

w m <Dmass k [ ] -1 d EA 1 m w EA

-(Vc )

=

<D

=

A·

- + -

.(c

-

C ) dt mass k k EA w

volume of drop let [m3]

surface area of droplet [m2]

concentration [kg/m3]

refers to ethyl acetate phase refers to water phase

equilibrium constant (=2.8)

mass flux [kg/m]

diffusion constant [mis]

[ 4-46]

(39)

The diffusion constants can be calculated with:

where: D d Sh k= Sh D d diffusion coefficient (10-9 m2/s) diameter of droplet (0.003 m)

Sherwood, under this circumstances given by: Sh=2.0+0.66Reo.5·Sc°.33.

Sh in the droplet is 6.6.

Sh outside the drop let is 58.4.

[4-47]

When the differential equation (4-46) is solved with the boundairies t=O CE~O, the time to reach 70% equilibrium can be found by substituting CEA by (0.7 cW/m). The results are given in

table 24.

Table 24: Calculated values for equilibrium time calculation

t70% 190 s Sc 1000 Re 76.4 Sh k (mis) inside droplet (w) 6.6 2.2.10-6

outside drop let (EA)

58.4 2.0.10-5

The velocity of the falling water droplets

The velo city ofthe falling water drop lets can be calculated from the following force balance:

7t 3 ₍ ₎ _7t 2 ₂

0= -d g PEA - Pw

+

Cd -d PEA V drop

6 4

where d diameter of drop let [m]

P density PEA=800 [kglm3], Pw=998 [kglm3]

g gravity constant 9.81 [mls2]

Vdrop falling velocity of droplet [mis]

Cd dragg coefficient, under this circumstances given by:

150 Cd

=2.3+-Re The falling velocity can be calculated by iteration.

The final dragg coefficient is 4.26.

The final falling velocity ofthe droplets is 0.048 mis.

REMARK: This is the falling velo city relative to the extracting fluid, not relative to the column.

[4-48]

(40)

The column height, the column diameter and the falling velocity relative to the column can be calculated, with the flows through the column known, by solving the following equations:

h

=

V drop' N· t₀₇ [4-50] where: h N tO.7 D D

=

~

4 V tot [ 4-51 ] n·h <1> v

=

1 - v

=

1 _ V,EA drop EA N. t 07 .0.25nD 2 height of column [m] number of real stages [-]

time to reach 70% equilibrium [sJ diameter of column [m]

volume holdup in column [m3], given by:

<1> <1> V V,EA V,W

tot N. t

0.7

VEA velocity of ethyl acetate relative to column [mis]

<1>v volume flow rate [m3/s]

[ 4-52]

[4-53]

Looking at the volume flow rate through the column, consider that in the first two stages (at the top) the flow changes a lot by the transport of most of the TMP at those two stages. At stage 3 to 13 the volume flow rate is nearly constant. Therefore the top stages need a larger diameter than the other stages. These diameters are ca1culated apart from the other stages. The diameters and heigths are given in table 25.

Table 25:Dimensions of the column.

stage diameter (m) height (m)

1 0.59 2.30 2 3 to 13 0.43 0.38 2.50 28.57

The total height ofthe column can be calculated by:

Hi is the height of stage i [m]

For liquid removal and liquid injection, 2 m auxiliary column height is necessary at both column ends.

The total height ofthe column is estimated at 37.4 m.

[4-54]

(41)

A critical note can be placed by the fact that in the first two stages about 95% of all the TMP is extracted. It is expectable that TMP is less effective extracted in these stages, so that in more stages on the top the diameter has to be larger than calculated. To find out what exactly is happening, experiments with a pilot tower are strongly recornrnended.

4.7. Crystalliser design

The final step in the purification of TMP from the reaction mixture is the crystallisation of the product from ethyl acetate. This is done in a CDC (Cooling Disc Crystalliser).

4.7.1. The Cooling Disc Crystalliser

Our choice to use the CDe for the crystallisation has been based on the following considerations:

The CDe consists of a horizontal trough which is divided by fixed cooling elements into separate compartiments, which are comparable to separate crystallisers. (Fig.5). In each compartiment a disc rotates which is equipped with teflon wip ers and mixing blades. The discs are mounted on a longitudinal axis which rotates at 5 to 45 rpm. Internal slurry transport facilities are absent in a eDe since both crystal and mother liquor flow freely from one crystalliser to the next through holes at the bottom of the compartiments and may amount up to 45 vol% in the last compartiment. Since the temperature difference between the slurry and the coolant is kept approximately constant throughout the eDe and the specific crystal surf ace area increases from inlet to outlet, the crystal growth rates are lowest in the last compartiments where the impurity concentration in the mother liquor is highe st. Consequently, in a eDe a good selectivity can be maintained at high production rates whereas the minimal slurry handling contributes to preserving a good quality.

In our design we took into account that the large st CDC available (Gouda machine factory) is 8 m long, has a volume of20 m3, contains 19 cooling elements with a total surface of 130 m2

and has a maximum capacity of about 30,000 tonnes solid product per year. 4.7.2 Design

The main problem here, is the cooling capacity ofthe crystalliser. The process fluid (48) has a temperature of 85 °C. A part ofthe cleaned liquid is recyc1ed (59) to decrease the inlet

temperature to 65°C. A smaller fraction of TMP in the solution leads to more secundairy nucleation in stead of primary nucleation. This means that larger crystals are formed. At a temperature of 30°C the liquid is just saturated with the byproducts, so the crystallisation is stopped at this temperature to give large quantities of practicaly pure TMP. The byproducts stay in the solution.

The calculation ofthe total heat duty consist oftwo parts. First, the cooling ofthe liquid mixture from 65°e to 30oe, and second, the heat of crystallisation ofthe TMP.

where: Xc

.1cH

mass fraction crystal in slurry [kg/kg] heat of crystallisation [kJ/kg]

(42)

The cooling capacity in a

eDe

is about 0.5 to 1.5 kW/m2 . This gives the required cooling

area.

A.:ool

=

387 m2

To calculate the specific cooling area, A.:ooJ, consider one crystalliser compartiment between

two cooling elements. (fig.5)

D

~---:---~ Z

..

Fig. 5: Crystalliser compartiment

The volume ofthis compartiment is:

The cooling area is (there is no shell cooling)

Thus, the specific cooling area is:

l.nD2 A cool 4

=

V

t·

nD2z 2 [ 4-56] z

This means that the specific cooling area of the crystalliser only dependends on the distance between the compartiments and not on the crystaliser length or diameter.

When the total cooling area and the specific cooling area are known, the total crystal1iser

volume can be calculated.

Table 26: C'1:.stalliser dimensions

Earameter value dimension

N 3 D 2.25 m L 8 m V 20 m3 z 0.30 m A.:ool 130 m 2 Q 122.6 kW Qtotal 368 kW 4 - 25