SECTION 3: CLARIFICATION AND UTILITIES (1)

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S

ECTION

3:

C

LARIFICATION AND

U

TILITIES

(1)

Aaliyah Hoosenally

10428141

Alcoholics

Anonymous

(2)

i

1. E

QUIPMENT

S

PECIFICATION

C

ALCULATIONS ...1

1.1. Cold Stabilisation Heat Exchanger HX-02 ... 2

1.2. Skin Separation/Plate and Frame Filter FLTR-01 ... 12

1.3. Membrane Filter FLTR-02 ... 14

1.4. Water Filter FLTR-03 ... 15

1.5. Aging in Barrels BRL-01 ... 16

1.6. Boiler HTR-01 ... 18

2. D

ETAILED

D

ESIGN

C

ALCULATIONS ... 19

2.1. Basic Heat Exchanger Design ... 20

2.2. Tubesheet Detailed Design ... 20

2.3. End Plate Design ... 23

2.4. Vessel Support Design ... 24

List of Tables

Table 1: Racked Red Wine Physical Properties ... 2

Table 2: Stabilised Red Wine Physical Properties ... 2

Table 3: Standard Tube Lengths and Outer Diameters ... 4

Table 4: Tube Length and Shell Diameter Selection 1 ... 5

Table 5: Tube Length and Shell Diameter Selection 2 ... 7

Table 6: Racked White Wine Physical Properties ... 10

Table 7: Stabilised White Wine Physical Properties ... 10

Table 8: Skin Separation/Plate and Frame Filter FLTR-01 Properties ... 12

Table 9: Membrane Filter FLTR-02 Properties ... 14

Table 10: Water Filter FLTR-03 Properties ... 15

Table 11: Physical Properties of Water for Hot Caustic Solution ... 18

Table 12: Fixed Tubesheet Effective Pressure Calculation ... 21

Table 13: Shell Longitudinal Stress Calculation ... 22

Table 14: Tube Longitudinal Stress Calculation ... 22

Table 15: Tube-to-Tubesheet Joint Loads ... 23

Table 16: Flanged Plate Ends - Bolted Covered with Full Faced Gasket ... 23

Table 17: Bolt Spacing ... 24

Table 18: Shell Mass Calculation ... 25

Table 19: Tube Mass Calculation ... 25

Table 20: Liquid Mass Calculation ... 25

Table 21: Flange End Plate (Flat Ends) Mass Calculation ... 25

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1

1.

E

QUIPMENT

S

PECIFICATION

C

ALCULATIONS

Aaliyah Hoosenally

10428141

Alcoholics

Anonymous

(4)

2

1.1. Cold Stabilisation Heat Exchanger HX-02

As mentioned in the equipment design section, this heat exchanger has been designed specifically to cool the red racked wine from 25°C to 15°C.

Table 1: Racked Red Wine Physical Properties

Racked Red Wine Inlet Outlet Mean

Temperature (°C) 25 15 20

Cp (kJ/kg.°C) 4.375 4.325 4.35

k (W/m.°C) 0.515 0.505 0.51

ρ (kg/m³) 973.5 976.5 975

μ (mPa.s) 1.8 2.3 2.05

mass flow rate (kg/s) 1.2188

Table 2: Stabilised Red Wine Physical Properties

Stabilised Red Wine Inlet Outlet Mean

Temperature (°C) 0 10.1235 5.062

Cp (kJ/kg.°C) 4.25 4.3 4.275

k (W/m.°C) 0.49 0.5 0.495

ρ (kg/m³) 982 978 980

μ (mPa.s) 3.8 2.6 3.2

(Table 5.2, p. 92, Rankine, 2004)

Using the specified temperature drop for the Racked Wine stream, the total amount of heat transferred can then be calculated. This can then be used to find the outlet temperature of the Stabilised Wine:

Total amount of heat transferred, Q =mCp∆T (kW) 53.0156 Temperature Out of Stabilised Wine To=Q/(mCp) + Ti(°C) 10.1235 An Overall Heat Transfer Coefficient must now be estimated:

Overall Heat Transfer Coefficient, Uass (W/m².C) 100

(Table 12.1, p. 637, Coulson, Richardson & Sinnot, 2005: Organic Solvent to Organic Solvent: 100 - 300 W/m².K)

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3 The log mean temperature must also be calculated using the following equation:

dT1 (C) 10

dT2 (C) 10.124

dT2 - dT1 (C) 0.124

dT2 / dT1 (C) 1.012

(∆T)lm (C) 10.062

Racked Wine will flow through the tubes, and the Stabilised Wine will flow through the shell so:

Tube In Temperature, t1 25

Tube Out Temperature, t2 15

Shell in Temperature, T1 0

Shell out Temperature, T2 10.1235

The correction factor, F1-2, can now be calculated using two dimensionless temperature ratios, R and S. If F1-2 is greater than 0.8, a shell and tube heat exchanger with 1 shell and any multiple of 2 tube passes is acceptable.

(Eq. 12.6 to 12.8, pp. 655-656, Coulson, Richardson & Sinnot, 2005)

R 1.0124

P 0.4

F1-2 0.9192

As F1-2 is greater than 0.8, we will begin with 1 shell with 4 tube passes – 4 passes because the volumetric flow of the wine is quite small, hence 2 passes would result in a tube side velocity that would be too low.

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4 The estimated heat transfer area can now be calculated:

Heat Transfer Area, A=Q/(U(∆T)lmF1-2) (m²) 57.3241

The table below displays the standard sizes of tube lengths and outer diameters (OD):

Table 3: Standard Tube Lengths and Outer Diameters

Length (ft) Length (m) Tube OD, do (in) Tube OD, do (m)

6 1.8288 0.250 0.00635 8 2.4384 0.375 0.009525 10 3.048 0.500 0.0127 12 3.6576 0.625 0.015875 16 4.8768 0.750 0.01905 20 6.096 1.000 0.0254 1.250 0.03175

If we use a standard tube OD of 0.0127m (1/2’’ in.) and a tube length of 1.8288m (6 feet), we can find the surface area of one tube, the total number of tubes required, the number of tubes required per pass, tube cross-sectional area, total area per pass, and the tube side velocity. The bundle diameter and shell clearance needed for each bundle can then also be calculated to find the optimal shell inner diameter.

A tube OD of 0.5’’ was chosen due to the small volumetric flowrate of the wine. Mechanical cleaning of the tubes is not necessary; hence the small tube diameter will not be a problem.

Tube OD do (m) 0.0127

Tube Length L (m) 1.8288

Number of Tube Passes 4

Tube Cross-Section Area Atc=π(do/2)2 (m²) 0.0001 Volumetric Flow Rate Q=ρ*mass flow (m³/s) 0.0013 Tube Surface Area At=πdoL (m²) 0.0730

Nt = A/At 786

Tubes per pass Np = Nt/4 196

Area per pass Ap = NpAtc (m²) 0.0249 Tube Side Velocity, ut = Q/Ap (m/s) 0.0502

K1 0.175

Table 12.4, p. 649, Coulson, Richardson & Sinnot, 2005

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5

n1 0.5037

Bundle Diameter Db = do(Nt/K1)1/n1 (m) 0.0130

Equation 12.3b, p. 648, Coulson, Richardson & Sinnot, 2005

Shell Clearance (m) 0.5167

Figure 12.10, p. 646, Coulson, Richardson & Sinnot, 2005

Shell ID Ds = Db + Shell Clearance (m) 3.5395

Optimum L/D ratio will fall between 5-10, p. 645, Coulson, Richardson & Sinnot, 2005

The same calculation can be performed for all the different tube lengths to find which length will give the optimum Shell ID:

Table 4: Tube Length and Shell Diameter Selection 1

Lt (m) At (m²) Nt Np Ap (m²) ut (m/s) Db (m) Lc (m) Ds (m) L/D 1.8288 0.0730 786 196 0.0249 0.0502 0.5037 0.0130 0.5167 3.5395 2.4384 0.0973 589 147 0.0187 0.0670 0.4441 0.0124 0.4565 5.3414 3.048 0.1216 471 118 0.0149 0.0837 0.4028 0.0120 0.4148 7.3484 3.6576 0.1459 393 98 0.0124 0.1005 0.3719 0.0117 0.3836 9.5352 4.8768 0.1946 295 74 0.0093 0.1340 0.3279 0.0113 0.3392 14.3792 6.096 0.2432 236 59 0.0075 0.1675 0.2974 0.0110 0.3083 19.7700

The above table shows that the tube length of 3.66 m that seems to give the optimum shell ID. The tube side and shell side heat transfer coefficients can now be calculated, in order to check whether the initial estimate of the overall heat transfer coefficient is correct.

Tube Side HTC Tube ID di (m) 0.0094 Re = ρutdi/μ 449.1297 Pr = Cpμ/k 17.4853 L/di 389 jh 0.0043

Using Figure 12.23, p. 665, Coulson, Richardson & Sinnot, 2005

Nu = jhRePr1/3 5.0126

ht =Nuk/di (W/m².C) 272.0190

To calculate the shell side heat transfer coefficients, the following formulas can be used to find the cross-sectional area of the shell, As, and the equivalent diameter, De, for a square pitch arrangement:

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6 (Equations 12.21 and 12.22, p. 672, Coulson, Richardson & Sinnot, 2005)

Shell Side HTC

Tube Pitch pt = 1.25do 0.0159 Baffle Spacing, lB =0.2Ds (m) 0.0767

p. 652, Coulson, Richardson & Sinnot, 2005

Cross Sectional Area, As (m²) 0.0059 Equivalent Diameter, De (m) 0.0090 Volumetric Flow Rate (m³/s) 0.0013 Shell Side Velocity, us (m/s) 0.2124

Re = ρusDe/μ 586.5296 Pr = Cpμ/k 27.6364 Baffle Cut % 25 jh 0.0220 Nu = jhRePr1/3 39.0127 hs =Nuk/De (W/m².C) 2141.5035

The overall heat transfer coefficient can now be calculated using the following equation:

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7

Overall Coefficient

Tube kw (W/mC ) 16 Stainless Steel 316 Tube

hod (W/m².C) 3000 Fouling Factors taken from _{'Organic Liquids'} Table 12.2, p. 640, Coulson, Richardson & Sinnot, 2005

hid (W/m².C) 3000

1/Uo (m².C/W) 0.0063

Uo (W/m².C) 157.7759

The error in the overall heat transfer coefficient that was used initially can now be found: Error in Uo = Uo-Uass/Uo 58%

This is an unacceptable error; hence we will have to make some adjustments:

Since the tube and shell side velocities are so small, we will have to increase the number of tube passes from 4 to 8 passes.

This will increase the heat transfer coefficient; hence will use a new assumed overall heat transfer coefficient of 300 W/m².C.

New Overall Heat Transfer Coefficient, Uass (W/m².K) 300.0000

New Heat Transfer Area, A (m²) 19.1080

Number of Tube Passes 8

K1 0.0365 Table 12.4, p. 649,

Coulson, Richardson & Sinnot, 2005

n1 2.675

Table 5: Tube Length and Shell Diameter Selection 2

Lt (m) At (m²) Nt Np Ap (m²) ut (m/s) Db (m) Lc (m) Ds (m) L/D 1.83 0.0730 262 33 0.0041 0.3016 0.3508 0.0115 0.3624 5.0503 2.4384 0.0973 196 25 0.0031 0.4019 0.3173 0.0112 0.3285 7.4236 3.048 0.1216 157 20 0.0025 0.5024 0.2899 0.0109 0.3008 10.1320 3.6576 0.1459 131 16 0.0021 0.6029 0.2708 0.0107 0.2815 12.9917 4.8768 0.1946 98 12 0.0016 0.8039 0.2432 0.0104 0.2536 19.2270 6.096 0.2432 79 10 0.0012 1.0048 0.2237 0.0102 0.2340 26.0530

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8 Tube Side HTC Tube ID di (m) 0.0094 Re = ρutdi/μ 1796.5189 Pr = Cpμ/k 17.4853 L/di 259 jh 0.002 Nu = jhRePr1/3 9.3258 ht =Nuk/di (W/m².C) 506.0819 Shell Side HTC Tube Pitch pt = 1.25do 0.0159 Baffle Spacing, lB =0.2Ds (m) 0.0657 Cross Sectional Area, As (m²) 0.0043 Equivalent Diameter, De (m) 0.0090 Volumetric Flow Rate (m³/s) 0.0013 Shell Side Velocity, us (m/s) 0.2896

Re = ρusDe/μ 799.8974 Pr = Cpμ/k 27.6364 Baffle Cut % 25 jh 0.0185 Nu = jhRePr1/3 _44.7404 hs =Nuk/De (W/m².C) 2455.9085 Overall Coefficient Tube kw (W/m.C ) 16 hod (W/m².C) 3000 hid (W/m².C) 3000 1/Uo (m².C/W) 0.0040 Uo (W/m².C) 251.2128 Error in Uo = Uo-Uass/Uo -16%

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9 This error is acceptable so we can now check the tube and shell side pressure drops:

The equation for the tube side pressure drop is:

(Equation 12.20, p. 667, Coulson, Richardson & Sinnot, 2005). For the shell side pressure drop:

(Equation 12.26, p. 675, Coulson, Richardson & Sinnot, 2005).

jf (tube) 0.0078

Figure 12.24, p.668, Coulson, Richardson & Sinnot, 2005

∆Pt (kPa) 22.3833

jf (shell) 0.075

Figure 12.30, p. 674, Coulson, Richardson & Sinnot, 2005

∆Ps (kPa) 38.1766

Both these pressure drops are low but acceptable if the nozzles are assumed to have a further pressure drop of 10kPa. The tube OD could be decreased further, or the number of passes increased further, to increase fluid velocity, but this would make it extremely difficult to clean the tubes thoroughly or to fit the tube passes into a fixed tube exchanger. Ease of cleaning is important in this winery, due to the frequent changes between grape types.

The nozzle and flange size can now be selected for the heat exchanger by estimating the optimal diameter of a nozzle for each inlet and outlet using the following equation:

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10

Flanges & Nozzles Tube Inlet Tube Outlet Shell Inlet Shell Outlet

Mass Flowrate, G (kg/s) 1.2188 1.2188 1.2250 1.2250

Density, ρ (kg/m³) 973.5 976.5 982 978

Optimal Diameter, Dopt (mm) 31.6639 31.6347 31.6623 31.7011

Nominal Diameter, DN (mm) 32 32 32 32

Nozzle Outer Diameter (mm) 42.2 42.2 42.2 42.2

Inner Diameter (mm) 36.66 36.66 36.66 36.66

Flange Size 32 32 32 32

Flange OD (mm) 120 120 120 120

(Nozzle – Schedule 10)

White Wine:

Using the overall heat transfer coefficient and heat transfer area of the heat exchanger that has been design to cool the racked red wine from 25°C to 15°C, it is possible to find out what the output temperature of the shell and tube fluid will be when the heat exchanger is used with white wine: Physical Properties:

Table 6: Racked White Wine Physical Properties

Racked White Wine Inlet Outlet Mean

Temperature (°C) 15 8.9 11.9

Cp (kJ/kg.°C) 4.325 4.275 4.3

k (W/m.°C) 0.505 0.495 0.5

ρ (kg/m³) 976.5 980 978.25

μ (mPa.s) 2.3 3.2 2.75

Table 7: Stabilised White Wine Physical Properties

Stabilised White Wine Inlet Outlet Mean

Temperature (°C) 0 6.1044 3.102

Cp (kJ/kg.°C) 4.25 4.3 4.275

k (W/m.°C) 0.49 0.5 0.495

ρ (kg/m³) 982 978 980

μ (mPa.s) 3.8 2.6 3.2

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11 It can be seen from the above tables that the racked white wine exits the heat exchanger at 8.9°C while the stabilised white wine is heated up to 6.1°C. This was calculated using the solver function on excel, solving for the racked white wine outlet temperature which gave the previously calculated final heat exchange area, using the final calculated overall heat transfer coefficient.

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12

1.2. Skin Separation/Plate and Frame Filter FLTR-01

The Plate and Frame filter is used in two separate processes in the winery. It is first used as a ‘Skin Separation Filter’ for white wine, to remove any remaining (large) solids from the free run and press fraction before fermentation. It is then reused later on in the process to re-filter the wine before it is membrane filtered. At this stage (clarification), any remaining particles larger than 0.3μm are removed.

The following information was taken from http://www.stpats.com/index.htm. This filter has been designed based on the KAPPA-5 40x40 Plate and Frame Filter.

Table 8: Skin Separation/Plate and Frame Filter FLTR-01 Properties

Process Skin Separation Clarification

Recommended Flow Rate (litres/hour.m²) 1500 500

Differential Pressure (kPa) 202.65 151.9875

Minimum Gauge Exit Pressure Required (kPa) 30.3975 30.3975

Minimum Required Gauge Inlet Pressure (kPa) 233.0475 182.385

Filter Sheet size 40x40 40x40

Filter Sheet Area (m2) 0.16 0.16

Maximum Number of Plates 40 40

Maximum Total Filter Area (m²) 6.4 6.4

Maximum Actual Flow Rate (litres/hour) 9600 3200

Approx. Total Volume to be Filtered (litres) 20000 16000

Total Time to be Filtered (hours) 2.08 5.00

Example of Calculations for Skin Separation Filter:

Minimum Required Gauge Inlet Pressure (kPa) = Differential Pressure + Minimum Gauge Exit

Pressure Required = 202.65 + 222.915 =233.0475

Maximum Total Filter Area (m2) = Filter Sheet Area (m2) x Maximum Number of Plates = 0.16 x 40 = 6.4

Maximum Flow Rate (litres/hour) = Recommended Flow Rate (litres/hour.m²) x Maximum Total

Filter Area (m2) = 1500 x 6.4 = 9600

Total Time to be Filtered (hours) = Approx. Total Volume to be Filtered (litres) / Maximum Flow

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13

Clarification:

Minimum Required Gauge Inlet Pressure (kPa) =182.385 Maximum Total Filter Area (m2) = 6.4

Maximum Flow Rate (litres/hour) = 3200 Total Time to be Filtered (hours) = 5.00

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14

1.3. Membrane Filter FLTR-02

Membrane filters are used as the last step of clarification to ensure microbial stability. The filter size of the Absolute PES Cartridge Filter that has been chosen to membrane filter the wine before bottling is PES 0.45, as it has a small enough pore size to remove (nearly) all yeast and bacteria from wine – it has 99.999% efficiency.

The following information was taken from http://www.stpats.com/index.htm.

Table 9: Membrane Filter FLTR-02 Properties

Membrane Filter type Absolute PES Cartridge Filter

Pore Size (microns) 0.45

Flow for 10" cartridge (0.7m²) (GPM/psid) 4.1

Flow for 10" cartridge (0.7m²) (litres per minute / kPad) 2.2510

Wine/Beer Maximum Operating Differential Pressure (psid) 30

Wine Maximum Operating Differential Pressure (kPad) 206.8427

pH 1 to 13

Maximum Working temperature (°F) 150

Maximum Flow Rate (GPM) at 30 psid 123

Maximum Flow Rate (litres/minute) at 207 kPad 465.61

Actual Approximate Flow Rate out of Ageing (litres/minute) 450

Volume to be Filtered per type of wine (litres) 35000

Total Time for Membrane Filtration (hours) 77.78

Maximum Flow for 10" cartridge (0.7m²) (litres per minute / kPad) = Maximum Flow for 10"

cartridge (0.7m²) (GPM/psid) x 0.549027562 (psid/GPM x litres/minute) = 2.25

Wine Maximum Operating Differential Pressure (kPad) = Wine/Beer Maximum Operating

Differential Pressure (psid) x 6.894757 (kpa/psi) = 206.84

Maximum Flow Rate (GPM) = Maximum Flow for 10" cartridge (0.7m²) (GPM/psid) x Wine/Beer

Maximum Operating Differential Pressure (psid) = 123

Maximum Flow Rate (litres/minute) = Maximum Flow for 10" cartridge (0.7m²) (litres per minute /

kPad) x Wine Maximum Operating Differential Pressure (kPad) = 465.61

Total Time for Membrane Filtration (hours) = Volume to be Filtered per type of wine (litres) / Actual

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15

1.4. Water Filter FLTR-03

The Absolute PES Cartridge Filter with a PES 0.22 pore size will be utilised to produce sterile water, which will be used for the final cleaning rinse over equipment.

The following information was taken from http://www.stpats.com/index.htm

Table 10: Water Filter FLTR-03 Properties

Membrane Filter type Absolute PES Cartridge Filter

Pore Size (microns) 0.22

Flow for 10" cartridge (0.7m²) (GPM/psid) 3.2

Flow for 10" cartridge (0.7m²) (litres per minute / kPad) 1.7569

Wine/Beer Maximum Operating Differential Pressure (psid) 30

Wine Maximum Operating Differential Pressure (kPad) 206.8427

pH 2 to 13

Maximum Working temperature (°F) 150

Maximum Flow Rate (GPM) at 30 psid 96

Maximum Flow Rate (litres/minute) at 207 kPad 363.40

Actual Approximate Flow Rate (litres/minute)

Volume to be Filtered (litres) 20000

Total Time for Membrane Filtration (hours) 0.92

Maximum Flow for 10" cartridge (0.7m²) (litres per minute / kPad) = Maximum Flow for 10"

cartridge (0.7m²) (GPM/psid) x 0.549027562(psid/GPM x litres/minute) = 1.76

Wine Maximum Operating Differential Pressure (kPad) = Wine/Beer Maximum Operating

Differential Pressure (psid) x 6.894757 (kPa/psi) = 206.84

Maximum Flow Rate (GPM) = Maximum Flow for 10" cartridge (0.7m²) (GPM/psid) x Wine/Beer

Maximum Operating Differential Pressure (psid) = 96

Maximum Flow Rate (litres/minute) = Maximum Flow for 10" cartridge (0.7m²) (litres per minute /

kPad) x Wine Maximum Operating Differential Pressure (kPad) = 363.4

Total Time for Membrane Filtration (hours) = Volume to be Filtered (litres) / Actual Approximate

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16

1.5. Aging in Barrels BRL-01

An example calculation will be shown in detail for Chardonnay:

Total grapes arriving (tonnes) = 80

Wine to Ageing (tones/5 tonnes grapes input into Crusher Destemmer) = 2.1324

Note. See Stream and Mass Balance Table in Volume 2 Supplement, PFD Location: CHPR4401-A-0102.

% Wine to Ageing = Wine to Ageing (tones/5 tonnes grapes input into Crusher Destemmer) / 5

(tonnes) x 100 = 43

Total Wine at Ageing (tonnes) = % Wine to Ageing x Total grapes arriving (tonnes) / 100 = 34.12 Temperature of Wine at Ageing = 20°C

Density of Wine at Ageing (kg/m³) = 975 (Table 5.2, p. 92, Rankine)

Total Wine at Ageing (litres) = Total Wine at Ageing (tonnes) x 1000 (kg/tonne) x 1000 (litres/m³) /

Density of Wine at Ageing (kg/m³) = 34993.23

Bordeaux Barrel Volume (litres) = 225

Number of Barrels Required = Total Wine at Ageing (litres) / Bordeaux Barrel Volume (litres) =

155.53 ~ 156 barrels required

Sauvignon Blanc:

Total Grapes (tonnes) 80

Wine to Ageing (tonnes/5 tonnes grapes) 2.1324

% Wine into Ageing 43%

Total Wine (tonnes) 34.1184

Density (kg/m3) 975

Total Wine (litres) 34993.23

Barrel Volume (litres) 225

Number of Barrels Required 156

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17

Shiraz:

Red Wine to Ageing (tonnes/5 tonnes grapes) 2.5532

% Red Wine to Ageing 51%

Total Red Wine (tonnes) 33.1916

Density (kg/m3) 975

Total Wine (litres) 34042.67

Cabernet Sauvignon:

Red Wine to Ageing (tonnes/5 tonnes grapes) 2.5532

% Red Wine to Ageing 51%

Total Red Wine (tonnes) 61.2768

Density (kg/m3) 975

Total Wine (litres) 62848

This gives a total of 430 barrels of red wine.

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18

1.6. Boiler HTR-01

The boiler will be used to heat water in order to produce a 5% hot caustic solution that is 80°C for cleaning. To find a suitable boiler, we must first find the power required to do this:

Water Flow Rate (L/minute) = 100

Water Flow Rate (m³/s) = Water Flow Rate (L/minute) / 1000 (L/m³) / 60 (seconds/minute) = 0.00167

Table 11: Physical Properties of Water for Hot Caustic Solution

Boiler In Out Mean

Temperature (°C) 20 80 50

ρ (kg/m3) 998.3 972 985.15

Cp (kJ/kg.K) 4.183 4.198 4.1905

Mass Flow Rate In (kg/s) = Water Flow Rate (m³/s) x ρin (kg/m3) = 1.664

Required Duty (kW) = Mass Flow Rate In (kg/s) x Cpmean (kJ/kg.K) x (Tout (°C) – Tin (°C)) = 418.34

Available sizes include: 30, 60 120, 160 and 500 kW (TCG Boiler Hire Brochure, http://www.tcgboilerhire.com.au/Downloads/TCG_Boiler_Hire_Brochure.pdf)

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19

2. D

ETAILED

D

ESIGN

C

ALCULATIONS

:

HX-01

Aaliyah Hoosenally

10428141

Alcoholics

Anonymous

(22)

20 See Section 1.1 for basic heat exchanger design.

2.2. Tubesheet Detailed Design

Using RCB-7.132, p. 39, TEMA, 1999, we can calculate the effective tubesheet thickness needed in order to resist bending stresses:

The mean ligament efficiency, η, is calculated using the following equation:

Stress in Tension, S (psi) 18709.87 Table 3.3.1(B), AS 1210, 1997

Shell ID, Ds (inches) 13.6236

Mean Ligament Efficiency, η 0.4195

Effective Design Pressure, P (psi) 20.3169 See later for calculation

Correction Factor, F 0.8 Figure RCB-7.132, p. 40, TEMA, 1999

Effective Tubesheet Thickness, T (inches) 0.1848

Minimum Tubesheet Thickness (inches) 0.3750 RCB-7.13, C-7.131, p. 38, TEMA, 1999

Actual Tubesheet Thickness (inches) 1

Actual Tubesheet Thickness (m) 0.0254

Tubesheet Type Supported As fixed tube exchanger

We must also check if shear stresses are controlling, in which case a different formula must be used to calculate effective tubesheet thickness. Shear will not control when:

(RCB-7.133, p. 43, TEMA, 1999).

P/S 0.001086

1.6 (1 - do/pitch)2 0.064

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21 expansion pressure, effective shell side design pressure, and effective tube side design pressure must first be calculated. Refer to RCB-7.161 to RCB-7.65, pp. 48-51, TEMA, 1999, for equations:

Table 12: Fixed Tubesheet Effective Pressure Calculation

J (assuming no expansion joint needed) 1 See later for checking this. Elastic Modulus of Shell Material, Es (psi) 2.87E+07 Table 3.3.7, p.84,

AS 1210, 1997 Elastic Modulus of Tube Material, Et (psi) 2.85E+07

Elastic Modulus of Tubesheet, E (psi) 2.86E+07 Shell Wall Thickness, ts (inches) 0.1882 Tube Wall Thickness, tt (inches) 0.0650 Differential Thermal Growth, ∆L (inches) 0

Tube and shell both constructed of Stainless Steel 316

Tube Length b/w outer faces, Lt (inches) 2.4384 Tube Length b/w inner faces, L (inches) 2.3876

Tubesheet Thickness Used, T (inches) 1

Shell OD, Do (inches) 14.0

Tube OD, do (inches) 0.5

Number of tubes in shell, N 200

Correction Factor, F 0.8

Figure RCB-7.132, p. 40, TEMA, 1999,

K 0.4624 RCB-7.161, p. 48, TEMA, 1999

Fq 4.0454 RCB-7.161, p. 48, TEMA, 1999, 1999

Differential Expansion Pressure, Pd (psi) 0 RCB-7.161, p. 48, TEMA, 1999, 1999 Equivalent Bolting Pressure- tube side, PBt (psi) 0 RCB-7.162, p. 49, TEMA, 1999, 1999 Equivalent Bolting Pressure- shell side, PBS (psi) 0 RCB-7.162, p. 49, TEMA, 1999

Shell Side Design Pressure, Ps (psi) 39.8854 RCB-7.163, p. 50, TEMA, 1999

fs 0.7306 RCB-7.163, p. 50, TEMA, 1999

Ps' (psi) 14.0690 RCB-7.163, p. 50, TEMA, 1999 Effective Shell Side Design Pressure (psi) 14.0690 RCB-7.163, p. 50, TEMA, 1999 Tube Side Design Pressure, Pt (psi) 39.8854 RCB-7.164, p. 51, TEMA, 1999

ft 0.9993 RCB-7.164, p. 51, TEMA, 1999

Pt' (psi) 20.3169 RCB-7.164, p. 51, TEMA, 1999 Effective Tube Side Design Pressure (psi) 20.3169 RCB-7.164, p. 51, TEMA, 1999 Fixed Tubesheet Effective Pressure, P (psi) 20.3169 RCB-7.165, p. 52, TEMA, 1999

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22 will be required to cope with any longitudinal stresses that might occur during the heat exchange. Refer to RCB-7.22 and RCB-7.23, pp. 53-54, TEMA, 1999, for equations:

Table 13: Shell Longitudinal Stress Calculation

Cs 0.5

RCB-7.22, p. 53, TEMA, 1999

Shell Wall Thickness, ts (inches) 0.1882

P1 = Pt - Pt' (psi) 19.5685

Ps' (psi) 14.0690

Ps* = P1 + Ps' (psi) 33.6374

Maximum Longitudinal Shell Stress, Ss (psi) 308.5953 Allowable Stress in Tension, S (psi) 18709.87

Table 14: Tube Longitudinal Stress Calculation

Fq 4.0454

RCB-7.23, p. 54, TEMA, 1999

Number of Tubes, N 200

Tube Wall Thickness, tt (inches) 0.0650

Tube OD, do (inches) 0.5

Ct 0.5

P2 10.4646

P3 6.8655

Pt* = P2 10.4646

Maximum Longitudinal Tube Stress, St (psi) 173.6772 Allowable Stress in Tension, S (psi) 18709.87

Hence, as both the shell and tube longitudinal stresses are well under the stress in tension of Stainless Steel 316 at the mean temperature, a shell expansion joint is not required.

The tube-to-tubesheet joint load must also be calculated in order to check whether the maximum load that may occur is allowable. See RCB-7.25, p. 56, TEMA, 1999 for equations:

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23

Fq 4.0454

RCB-7.25, p. 56, TEMA, 1999

Pt* = P2 10.4646

Number of Tubes, N 200

Maximum Effective Joint Load, Wj (lbf) 30.8548 Maximum Allowable Joint Load (lbf) 3673673

Hence the tube-to-tubesheet joint load is within an acceptable limit.

2.3. End Plate Design

The ends of the shell will be bolted flat end covers (blind flanges) with a full faced gasket. The minimum thickness required is given by:

(Equation 13.42, p. 818, Coulson, Richardson & Sinnot, 2005).

Table 16: Flanged Plate Ends - Bolted Covered with Full Faced Gasket

Design Constant, Cp 0.4

p. 818, Coulson,

Richardson & Sinnot, 2005

Bolt Circle Diameter, De (m) 0.470

ND=350, Figure v, Table D, AS 2129, 2000

Design Pressure (N/mm2₎ _0.275

Design Tensile Stress (N/mm2) 129 Required Plate Thickness (m) 0.0087

Actual Plate Thickness (m) 0.022

The bolt spacing of the flange must now be checked. The maximum bolt spacing for a flange with a full faced gasket is:

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24

Minimum Bolt Diameter (m) 0.0127

RCB-11.1, C-11.1, p.77, TEMA, 1999

Nominal Diameter 350

Figure v, Table D, AS 2129, 2000

Actual Bolt Diameter, dB (m) 0.0222

Flange Thickness, tf (m) 0.022 Gasket Type Flat Metal - Stainless Steel Gasket Factor 6.5 Table 3.21.6.4(A), p. 205, AS 1210, 1999

Elastic Modulus of Flange Material (Steel), E (MPa) 1.96E+05 Maximum Bolt Spacing, max Pb (m) 0.0632

Bolt Circle Diameter (m) 0.470 Figure v, Table D, AS 2129, 2000

Actual Number of Bolts 12

2.4. Vessel Support Design

The main sources of load to consider are 1. Pressure

2. Dead weight of vessel and contents 3. Wind

4. Earthquake 5. External Loads

The longitudinal stresses on the shell have already been shown to not be an issue in comparison to the tensile strength of the stainless steel. Wind loads would not be an issue either as the vessel is small and horizontal, and will most likely be places indoors. The chances of an earthquake occurring in Western Australia are extremely low; hence seismic loads do not have to be considered either. External loads will be to a minimum. Hence, the only significant load that must be considered is the dad weight of the vessel and contents: The shell and tubes will be constructed out of stainless steel pipe, hence their volume (steel volume) can be calculated by:

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25 Stainless Steel 316 Density (kg/m³) 8000

Actual Shell OD , Do (m) 0.3556

Actual Shell ID, Ds (m) 0.3460

Actual Shell Thickness (m) 0.0048

Shell Length, Ls (m) 3.556

Shell Volume (m³) 0.0187

Shell Mass (kg) 149.8697

Table 19: Tube Mass Calculation

Stainless Steel 316 Density (kg/m³) 8000

Tube OD (m) 0.0127 Tube Thickness (m) 0.0017 Tube ID (m) 0.0094 Tube Length (m) 2.4384 Tube Volume (m³) 0.00014 Tube Mass (kg) 1.1179 Number of Tubes 200

Total Tube Mass (kg) 223.5861

Table 20: Liquid Mass Calculation

Average Density (kg/m³) 977

Max Liquid Volume (m³) 0.3344

Max Liquid Mass (kg) 326.7374

The maximum liquid volume was calculated assuming the entire shell is filled with liquid.

Table 21: Flange End Plate (Flat Ends) Mass Calculation

Stainless Steel 316 Density (kg/m³) 8000 Blind Flange Diameter (m) 0.525 Blind Flange Thickness (m) 0.022

Blind Flange Volume (m³) 0.0048

Blind Flange Mass (kg) 38.0997

Number of Ends 2

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26

Shell Mass (kg) 149.8697

Max Liquid Mass (kg) 326.7374

Total End Mass (kg) 76.1993

Total Tube Mass (kg) 223.5861

Total Vessel Mass (kg) 776.3925 Sum of shell, liquid, tube and ends.

Correction factor 1.08

Taking into account any external fittings.

Corrected Vessel Mass (kg) 838.5039

Dead Weight of Vessel (kN) 8.225724 =9.81*Vessel Mass

Figure 13.26 (a), p. 847, Coulson, Richardson & Sinnot, 2005 can now be used to design saddle supports for the vessel (See Volume 2, Section 3, Chapter 3).