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Roger Adams Chemical 600 S. Mathews Ave. Urbana, IL 61801 DATE: September 30, 2009 TO: Victor Shum, Plant Supervisor FROM: Ben Mack, Melissa Medrano, Jason Pulley, Process Engineers SUBJECT: Double-Effect Evaporator Replacement Investigation Final Report REFERENCE: Memo Dated August 23, 2009, Double-effect Evaporator Replacement – Rotation 1 and 2; Double-Effect Evaporator Replacement Investigation Workplan 1. Abstract In order to fully understand the operation and effectiveness of the replacement double-effect evaporator, an investigation was carried out to determine the effects of steam and vacuum pressure on the heat transfer coefficients of each exchanger, the condensate flow rate at the end of the process, and the overall steam efficiency in the system. The steam pressure and 2 nd effect vacuum pressure were found to have a positive correlation with the 2 nd effect overhead condensate flow rate. Steam efficiency is to a large degree affected by changes in the steam and vacuum pressure, as an increase in steam pressure and vacuum pressure causes an increase in the steam efficiency. The two parameters have the same significance in regards to steam efficiency. The heat transfer coefficient for each evaporator heat was determined for different vacuum and steam pressures. The patterns indicate an increase in heat transfer coefficient for both effects with increasing steam pressure. Conversely, increasing the vacuum pressure causes a decrease in the heat 1

FINAL Evaporator Final Report

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Page 1: FINAL Evaporator Final Report

Roger Adams Chemical600 S. Mathews Ave.

Urbana, IL 61801

DATE: September 30, 2009

TO: Victor Shum, Plant Supervisor

FROM: Ben Mack, Melissa Medrano, Jason Pulley, Process Engineers

SUBJECT: Double-Effect Evaporator Replacement Investigation Final Report

REFERENCE: Memo Dated August 23, 2009, Double-effect Evaporator Replacement – Rotation 1 and 2; Double-Effect Evaporator Replacement Investigation Workplan

1. Abstract

In order to fully understand the operation and effectiveness of the replacement double-effect evaporator, an investigation was carried out to determine the effects of steam and vacuum pressure on the heat transfer coefficients of each exchanger, the condensate flow rate at the end of the process, and the overall steam efficiency in the system. The steam pressure and 2nd effect vacuum pressure were found to have a positive correlation with the 2nd effect overhead condensate flow rate. Steam efficiency is to a large degree affected by changes in the steam and vacuum pressure, as an increase in steam pressure and vacuum pressure causes an increase in the steam efficiency. The two parameters have the same significance in regards to steam efficiency. The heat transfer coefficient for each evaporator heat was determined for different vacuum and steam pressures. The patterns indicate an increase in heat transfer coefficient for both effects with increasing steam pressure. Conversely, increasing the vacuum pressure causes a decrease in the heat transfer coefficient and a decrease in the vacuum pressure causes an increase in the heat transfer coefficient for both effects. The heat transfer coefficient for the heat exchanger increases with increasing steam and vacuum pressure. However, the heat transfer coefficient for the heat exchanger decreases with decreasing steam and vacuum pressures. Literature has been found suggesting possible reasons for discrepancies in the heat transfer coefficients, including scaling, frothing, and bubbling of the liquid. The evaporator was predicted to run more effectively and efficiently under double-effect feed forward mode to ensure that the 90% minimum concentration would be met in the system. It was recommended for the Senior Chief Technician to use Careclean F Descalex P produced by Marine Care, as the cleaning product for the U.S. Coast Guard Ship scaled up evaporator due to its cleaning efficiency on heavy scale and safe use in evaporators. 2. Introduction

The suitability of replacing the current double-effect evaporator used for concentrating high fructose corn syrup by an old evaporator found in storage was investigated due to a scheduled

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shutdown for routine maintenance of the current double-effect evaporator. The shutdown will occur in approximately 8 weeks.1 Therefore, an immediate study of the old evaporator was necessary. The replacement of the evaporator was proposed due to current cost savings measures, specifically so that production is not temporarily halted during the routine maintenance period of the current double-effect evaporator. To simplify the process and to understand the effects of the different variables, this study was performed using industrial tap water instead of high fructose corn syrup as the non-concentrated solution and concentrated product. This study included an analysis of the steam efficiency of the replacement evaporator and the development of a model for predicting the replacement evaporator’s operation. The physical and economic feasibility of producing high fructose corn syrup using the replacement double-effect evaporator will also be determined in the second 4-week rotation, which is not included in this report.

This report will focus on the first 4 week term, Rotation 1 objectives, which included the calibration of the height of the sight glass level against volume for both effects, characterization of the double-effect evaporator for wide ranges of steam and vacuum pressures, the effects of steam and 2nd effect pressure on 2nd effect vapor condensate flow rate, and the effects of operating conditions on steam efficiency. In addition, the heat transfer coefficients in the evaporator effects were calculated from experimental data, and their changes as a function of all relevant parameters were established. Finally, single-effect and parallel-effect as well as double-effect feed backward modes were studied to determine the optimum effectiveness and efficiency in the double-effect evaporator. A recommendation for a product to clean a scaled up evaporator for a U.S. Coast Guard Ship is also provided.

The experimental procedure described in this report was based on the vacuum and steam pressures’ effects on the operation of the evaporation system. These measurements provided data on how each of these two variables alters the 2nd effect condensed vapor flow rate as well as the heat transfer coefficients. In addition, the most influential variables were examined and their effects determined.

3. Theory

The general purpose of evaporation is to boil off a volatile solvent (in this case water) from a nonvolatile solute to produce a more concentrated product solution.2 This concentrated product (often referred to as the liquor) is the stream of value in this process. In this case, the current double-effect evaporator is being used to concentrate a high fructose corn syrup from 10% to 90%. However, for experimental purposes, industrial water will be used in this rotation in place of the high fructose corn syrup. A multiple-effect evaporator uses a series of evaporators (called effects) to accomplish the previously stated goal. This system utilizes the latent heat from the water vapor from the previous effect to boil off excess water in the solution.3 Steam efficiency is greatly increased in a multiple-effect evaporation setup compared to a single-effect evaporator, as less direct steam is required to produce the same product concentration. This is due to the reuse of the water vapor boiled off the solution. Raw steam is used as the heating source for the 1st effect, and then the boiled off vapor from the 1st effect is used as the heating source for the 2nd

effect.4

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There are several different stream arrangements possible for a double-effect evaporator system used in industry: parallel-effect, forward-effect, and backward-effect. In a parallel-effect system, the feed stream is split and fed to each effect simultaneously. In a forward-effect, the feed stream is fed using a single pump, starting with the effect that the steam is directly fed to, and ending at the effect where concentrated product is collected. Figure 1 shows the general setup for a forward-feed double-effect evaporator similar to the one investigated on this rotation.

Figure 1: General Double-effect Evaporator Schematic3

The functionality of this equipment is based on heat-based separation principles. Solutions containing components with different boiling points can be separated through the application of heat, where the heat transfer from the steam line to the solution in the tank raises the temperature of the solution to its boiling point and (when enough heat is provided) vaporizes the more volatile component off of the solution. This heat transfer lowers the temperature of the steam, condensing it into a liquid.

The heat transfer for evaporators can be quantified using the following general equations (Refer to Appendix A for more rigorous calculation methods):2

q=UATlm (1)

q= Sλ (2)

Where q is the heat transfer rate, U is the heat transfer coefficient, A is the heat transfer area of the effect, T is the temperature difference between the steam into the evaporator and the boiling point of the solvent, S is the mass flow rate of the steam entering the evaporator, and is the latent heat of the steam. By setting these two equations equal to one another, it becomes possible to calculate the overall heat transfer coefficients (U) at different steam pressures to determine the steam pressure affect on the U value for the 1st and 2nd effects.

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The efficiency of the steam is defined as the amount of product produced divided by the amount of steam fed to the system. Multiple-effect systems increase the efficiency because the waste streams are used as heat sources rather than just direct fed steam.

Normally, evaporators are constructed from a form of steel, usually stainless steel to prevent damage from any ferrous metals or other corrosive materials that may be run through the system. One common problem in evaporation is scaling. Scaling occurs when buildup is deposited on the heating surfaces by the solution being boiled. This buildup has been reported to diminish the overall heat transfer coefficient. Normal solutions to this problem call for the shutdown of the evaporator to clean the heating surfaces.

The current system also makes use of the functionality of heat exchangers. Heat exchangers are designed with the purpose of transferring heat between streams of different energy. In some cases, heat exchangers utilize a heated vapor to raise the temperature of another stream or to vaporize the stream. In this case, heat exchangers are utilized as condensers after each effect to condense the waste steam into liquid water by passing a sufficient amount of cold water through the heat exchanger. Condensing the overhead vapor from the 2nd effect steam allows the vapor to be collected, quantified, and then discharged at atmospheric pressure. The main equations governing the heat exchangers are (also found in the calculations in Appendix A):2

q=UATlm (3)

q=mCpT+Hv (4)

Where q is the heat transferred, U is the overall heat transfer coefficient, A is the heat transfer area of the device, m is the mass flow rate, T is the temperature change of the stream, Cp is the heat capacity of the stream, and Tlm is the log mean temperature difference of the heat exchanger.

4. Experimental Apparatus and Procedure

The double-effect evaporator setup consists of two identical insulated effects, a heat exchanger used to condense the 2nd effect overhead, and two collection tanks used to collect the condensed overhead.6 Each of the effects was custom fabricated with an inner diameter of 11.75 inches and a height of 4 feet. The exchanger used to condense the 2nd effect vapor is an American Standard heat exchanger with a temperature rating of 450°F, tube pressure rating of 150psi, and a shell pressure rating of 225psi. The two effects also have small Blackmore and Giant brand heat exchangers that serve as condensers on the steam outlet lines. All of the exchangers used are of the shell and tube variety. More information about the effects and 2nd effect overhead condenser can be found in Table 1 below.

Table 1. Effect and exchanger characteristics6

1st Effect 2nd Effect 2nd Effect Vapor

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Condenser

Manufacturer Custom fabricated Custom fabricated American Standard

Feed (Hot Stream) RAL steam supply 1st effect overhead 2nd effect overhead

Feed (Cold Stream) RAL water supply RAL water supply RAL water supply

Water Flow Meter Brand

Neptune Hersey Hersey

Heat Transfer Area (ft2) 1.8 3.3 16

Low-pressure steam (approximately 35psi) enters the 1st effect from the RAL steam supply line and is fed through a tube bundle side mounted near the bottom of the 1st effect. The overhead stream of this effect is then used to boil water in the 2nd effect. Steam feed pressure is monitored using two analog pressure gauges: one before the steam flow throttling valve to measure supply pressure and one after the valve to measure the actual pressure of the steam added to the evaporation system. Each effect has a mounted analog pressure gauge to monitor vessel pressure. The vacuum pressure on the 2nd effect can be monitored and controlled off of the main vacuum line that is located on the south wall of RAL 8 with the process water lines.

The process water is metered from the building supply line and is fed to both of the effects. Unlike the 1st effect flow meter, the 2nd effect and vapor overhead condenser flow meters record the volume of water that enters the vessels instead of the flow rate. Analog temperature gauges are used to monitor the feed temperature of the water to the 1 st effect, the cooling water inlet and outlet temperatures to the 2nd effect vapor condenser, and also on the hot stream outlet from the condenser. Digital gauges monitor the temperatures of the water near the bottom of the effects where the steam coils enter. Figure 2 below diagrams the setup of the evaporator unit. Each measurement point for temperature, pressure, and flow rate is indicated with squares, circles, and triangles, respectively.

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Figure 2. Double-effect evaporator design setup7

In order to investigate the effects of steam and vacuum pressure, several measurements were recorded per trial to perform the necessary calculations. Starting from a completely drained system, the 1st and 2nd effects were filled with water; the 1st effect feed is metered by a level controller, and the 2nd effect flow meter reads the actual volume of water being added to the effect. Calibration curves were created for both of these effects by adding known amounts of water and measuring the resulting height from grade in each sight glass. The calibration curve in the 2nd effect was the most crucial since it was used to quantify exactly how much water boiled out in the overhead. Initial and final heights in the sight glass were recorded and plugged into the calibration curve to measure this amount of water. The actual calibration curves used in the experiment can be found in Section 5.1.

Once the steam supply line was opened to the system, it typically took a couple of minutes for the line to stabilize. Stabilization was determined as the point in which the pressure gauge readings stopped increasing and leveled out. This pressure reading was indicated in Figure 2 as P1. Additionally, the steam supply temperature was measured and represented as T2. After the steam had passed through the 1st effect, it passed through a condenser before passing to the drain. By collecting the condensed steam out of this condenser, the steam flow rate to the 1 st effect, F2, was measured by collecting the water over a specified period of time. In the same fashion, the flow rate of the steam generated in the 1st effect, F4, was measured.

The inlet temperatures and flows of the water supply to the 1 st and 2nd effects were measured as indicated by T1 and T5, F1 and F3, respectively. These temperatures were used to determine

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specific enthalpy values (hf) used in heat transfer calculations around each effect. The flows were used to evaluate the amount of water boiled in each effect. To monitor the progression of the boiling, the thermocouples represented by T3 and T6 were utilized. The thermocouples provided temperature readings of the water in each effect and indicated when boiling had occurred. Once these temperatures reached the boiling point of water at the operating pressure of each effect (measured as P2 and P3 by pressure gauges on top of the effects), it was assumed that full boiling was rampant and that vapor was being transferred to the next unit in the system. The operating pressure of the 2nd effect, P3, was dictated by P4, the vacuum pressure of the system. The selected vacuum pressure for the system was the operating pressure of the 2nd effect.

After the 1st effect water began to boil, monitoring the temperature gauge indicated that vapor was in fact being transferred to the 2nd effect, and T4 confirmed this thought. Without any vapor present, this thermometer would read room temperature. The same reasoning follows with T7, the temperature of the 2nd effect vapor overhead. As the 2nd effect water boiled, the overhead was condensed using the heat exchanger in the system. Another indication of boiling in the 2nd effect was the outlet temperature of the cooling water, T9, which began to increase from its inlet temperature, T8. The flow rate of this cooling water, F5, dictated how much the temperature increased: a higher flow rate resulted in less of a temperature increase. This is based on equation 5 in the Appendix which shows that with the same heat transfer, q, increasing m will result in a decrease of ∆T.

Once the overhead condensed, it was collected in the two collection tanks. On its way to the tanks, the overhead vapor passed a temperature gauge, T10, which measured the final temperature of the condensed vapor. The condensed vapor flow rate was determined in the same fashion that steam flow rates were measured in the 1st and 2nd effect. One of the collection tanks was selected for collection prior to testing. During a timed interval, the collection tanks were switched so that only the vapor produced during that time period was collected in the 2nd tank. At the end of the collection period, the valves were switched so that the 1 st tank was once again collecting the condensed overhead. After the 2nd tank was isolated, the vacuum was lifted from it and the vapor was collected through one of the drains. In the ideal situation, the sight glasses of the collection tanks could have been calibrated to the volume in the tanks; however, during normal operation, the fluctuations in the sight glass level were significant, and it was deemed more reliable to manually collect samples and measure the exact amount collected rather than using a calibration.

5. Results and Discussion

Exploration of the effects of steam and vacuum pressure on the overall system performance was performed using a partial 3 level factorial design for 2 factors.8 Ideal experimentation would include a full 3 level factorial design for 2 factors; however, time constraints and lengthy sample collection periods prevented all 9 of these samples from being carried out. Figure 3 shows the chosen experimental design and which experiments were executed.

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Figure 3. Experimental 2 level partial factorial design for 3 factors. X’s represent data points not collected.

The decision was made to run a replicate of the high steam and high vacuum pressure data point in favor of running another unique sample point such as the medium/medium sample. Both of these points are valuable, but, with only a couple samples, replicates were deemed to be a higher priority instead of a unique middle data point. Data analysis based on only a couple of samples is difficult as small errors in a sample point can cause severe inaccuracies in the data trends. Having a replicate of a data point provides some validation to the data whereas a middle point may just make the data more unclear; for example, if the middle point does not fit the trend of the other couple data points, it is unclear which point is inaccurate and what the actual trend should look like. A summary of the design matrix detailing the order that the experiments were run can be found in Table 2 below.

Table 2. Experimental design matrix9 X1 (steam P, psig) X2 (vacuum P, in Hg)

10 010 1030 1730 1724 1730 8

This experimental design matrix was varied from the completely randomized design matrix presented in the workplan in order to better fit samples into time periods where they could be carried out.9

5.1. Calibration

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Accurate volume measurements were obtained by first calibrating the sight glasses for the effects to known volumes and creating calibration curves from this data. For each effect, an initial height on the sight glass was marked, and a specific amount of water, measured with the supply line flow meters, was added to the effect. The new height was marked and measured, and the process was repeated for several data points. Using these points, the calibration curves seen in Figures 4 and 5 below were determined.

Figure 4. 1st effect sight glass calibration curve

Figure 5. 2nd effect sight glass calibration curve

Equations for the regression lines and R2 values can be found in Table 3 below.

Table 3. 1st and 2nd effect calibration curve equations

Effect Calibration Curve Equations R2 Values

1st Height = 1.82*(volume) + 14.19 0.9986

2nd Height = 2.18*(volume) + 24.46 0.9996

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R2 values in Table 3 indicate that these calibration curves can be used to accurately translate sight glass height to volume for each of the 2 effects. As discussed in Section 4, the sight glasses on the collection tanks were not successfully calibrated due to oscillations in the sight glass height.

5.2. Data Analysis

For the low level steam pressure samples, a pressure of 10 psig was chosen. As the steam pressure is directly correlated to the flow rate, decreasing the steam pressure below 10 psig would not boil water in the 1st effect in a timely manner. During experimentation, 10 psig steam was capable of boiling water in the 1st effect within 1.5 hours. The rate that steam was produced in this effect, however, was insufficient to boil water in the 2nd effect. During this sample, water in the 2nd effect only reached a temperature of 82.3°F, which was far from the boiling point (212°F at atmospheric pressure). Therefore, data could not be collected for the heat exchanger since no vapor was produced to be condensed. This point was thus excluded from further analysis since an overall steam efficiency could not be calculated. Similar to the 1st sample, the 2nd sample at 10 psig steam and 10 in Hg also could not be used in data analysis since the low pressure steam inlet only raised the 2nd effect temperature to 138°F. Qualitative comparison of these 2 samples indicates the benefits of operating the 2nd effect under a vacuum; the vacuum helped the 2nd effect reach a higher temperature while the steam to the 1st effect remained the same. A summary of the data collected for these 2 samples can be found in Table 4 below. Full data can be found in Appendix B.

Table 4. Data summary for the 1st and 2nd trials Trial # 1 2Vacuum P (in Hg) 0 10Steam P (psig) 10 10Steam T (°F) 254 254Effect 1 Initial T (°F) 79.4 73.6Effect 1 Vapor T (°F) 200 201Effect 2 Initial T (°F) 74.3 75.6Effect 2 Vapor Temp (°F) 82.3 138

Based on the findings discussed above, it was deemed impractical to operate the steam pressure at 10 psig and to operate the vacuum pressure at 0 in Hg. The other 3 samples involving these pressures were consequently omitted from the factorial design due to time constraints and the proven poor results. These samples are marked with an X in Figure 3 above. The 4 th sample with an X in Figure 3 is a medium/medium sample and was only excluded due to time constraints.

For the remaining 4 data points collected, steam was generated in the 2nd effect allowing data for the heat exchanger to be collected. A summary of the key data for these 4 trials can be found in Table 5 below with complete data located in Section B of the Appendix.

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Table 5. Key data summary for Trials 3-6 Trial # 3 4 5 6Vacuum Pressure (in Hg) 17 17 8 17Steam Pressure (psig) 30 30 30 25

U1 (Btu/min °F ft2) 10.77 10.11 14.66 8.41

U2 (Btu/min °F ft2) 2.76 2.63 2.96 2.50

Uhe (Btu/min °F ft2) 1.27 1.29 0.49 1.16

Condensate Flow Rate (ml/min) 265 274 200 160

Efficiency (%) 74.39 84.31 58.26 57.76

Table 5 shows the key calculated variables: heat transfer coefficients (U1, U2, and Uhe), condensate flow rates, and efficiencies for each trial. Trending the vacuum and steam pressures individually against each key variable resulted in data that produced no distinct correlation. Figures of these trends can be found in Appendix C. Since only 4 data points were collected, it was difficult to know whether a correlation existed or if an outside source was causing inconsistencies with the data. One potential source for this variation was in the data collection. Samples taken at the highest possible vacuum pressure experienced fluctuations in the vacuum pressure; the pressure would cycle from 17 in Hg down to 8 in Hg and vice versa. To account for these vacuum pressure changes, sample points were taken over complete cycles by starting at 8 in Hg and measuring the time until it returns back to 8 in Hg. A potential explanation for this cyclic action is discussed in detail in Section 5.3. Sampling while the vacuum pressure is varying is not ideal since some of the samples may have the vacuum pressure stay at the maximum for longer than others. This would result in more water boiling in the 2 nd effect and, therefore, a higher q value for the heat exchanger. Since the heat transfer coefficient is calculated from this q value, the coefficient may be artificially high as is the case in Trial #5. Heat transfer coefficients should remain constant regardless of steam and vacuum pressures under normal operation.10 If scaling is present within the effects, the heat transfer area will decrease. By not taking this scaling into account and using the unscaled exchanger area in the calculations, the heat transfer coefficient will be calculated artificially low.

Since both steam and vacuum pressures affect the system simultaneously, a true correlation between the heat transfer coefficients and the pressures takes them both into account simultaneously. To accomplish this, correlations were determined using the ANOVA regression tool of Microsoft Excel to trend how the pressures influence these values together. Linear regression equations were found and the coefficients and intercepts of these equations as well as the R2 value and standard error can be found in Table 6 below. Evident from the R2 values, these regression equations show a strong correlation between the combination of the two pressures and the heat transfer coefficients.

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Table 6. Regression equations for heat transfer coefficients and efficiency

Quantity U1 U2 Uhe

Condensate Flow Rate Efficiency

Steam Pressure Coefficient 0.31 0.03 0.02 18.16 0.04Vacuum Pressure Coefficient -0.49 -0.03 0.09 6.96 0.02Intercept 9.27 2.39 -0.81 -400.61 -0.75R2 0.97 0.87 1.00 0.98 0.99Standard Error 0.84 0.12 0.01 0.07 0.03

Efficiency in the system is affected in the same fashion as the heat transfer coefficient: both pressures affect it at the same time, and their joint effect must be considered. The relationship between the pressures and efficiency can be seen in the regression line described in Table 6 above. Higher steam pressures increase the amount of steam produced in the 1st and 2nd effects and, hence, will lead to a higher efficiency. Additionally, the higher the vacuum pressure, the easier it is for the steam produced in the 2nd effect to be transferred to the heat exchanger and collection tanks. Lower vacuum pressures may result in some of the steam condensing in the effect before it can leave. This will decrease the efficiency as more energy is required to boil this water again.

Condensate flow rates are also affected by both pressures. The regression for this correlation can also be found in Table 6 above. The condensate collection flow rates have both system and experimental error associated with them.  The system error is due to the head built up in one of the collection tank drainage lines.  This head would not drain from the line unless significant water built up to force it out.  This head of water led to some possible inaccuracies in the measurement of the condensate flow rate.  Experimentally, the setup of the collection tanks left certain areas of error.  The vacuum presented difficulties in pulling the water into the proper collection tanks, as well as draining the tanks.  The vacuum proved to be difficult to operate effectively, which could have potentially led to errors in the experimental collection values.

5.3. Experimental Sources of Error

Besides the aforementioned potential sources of variability in the calculations, other sources of error exist in system. Most importantly, inherent fluctuations in the steam and vacuum pressures from the steam and vacuum supply lines affect results. Ideal operation occurs at the maximum allowable working steam pressure while under the strongest possible vacuum. However, due to the fluctuations in each pressure, running each supply fully open would prevent this ideal state from being maintained at a steady state. Since maximum pressures of both supply lines cannot be controlled, it is therefore desirable to operate each system at a pressure slightly below the maximum so that steady state can be reached. For the steam supply line, this means operating at 30 psig instead of at the maximum of 35 psig. This ensures the most amount of heat can be added to the 1st effect under steady state conditions. The vacuum pressure should be run at approximately 8 in Hg. This pressure peaks around 17 in Hg, but it cycles from 8 in Hg to 18 in Hg while trying to operate with the system fully open.

Using the throttle on the system prevented these oscillations and allowed the 2nd effect to reach steady state. This was very important since the water boiled easier under a higher vacuum

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pressure. If the pressure varied, the water boiled in one instant and, after a pressure drop, ceased boiling. This created cycles of periods where the collection tanks were collecting condensed overhead for a period of time and then not collecting after the pressure has dropped. In data collection, this led to some inconsistent results since timed samples collected during periods of pressure swings would lead to inaccuracies. Further investigation needs to be performed to identify the maximum allowable steady state working pressure. 8 in Hg was the chosen sample as typical vacuum pressure swings with the valve fully open settled at this pressure. As mentioned before, the vacuum pressure has significant effect on the operation of the 2nd effect, so it is recommended that the vacuum pressure be optimized to its highest possible level.

Another issue encountered with the vacuum system lies in the manual valve that controls the vacuum to the 2nd effect. In initial trials, this valve was opened completely and the main vacuum line valves were throttled. Operating in this mode led to some of the vapor from the 2nd effect flowing back through the vacuum line. Evidence for this backflow was that the vacuum line became very hot during operation. Loss of steam through the vacuum resulted in lost heat and affected heat transfer calculations since not all of the steam created will pass through the heat exchanger. Opening this valve only a fraction of the way in later trials limited the amount of vapor flowing through the vacuum system resulting in less steam loss and more accurate results.

Levels in the sight glasses on the 2nd effect and collection tanks were affected by the vacuum pressure and boiling process making accurate readings difficult to obtain. During boiling and collection, the levels in the sight glasses did not demonstrate accurate levels as air bubbles would form in them and levels would move up and down in the sight glass. In the 2 nd effect, initial readings were thus taken and compared to the final level once the effect stopped boiling and the water settled. This allowed overall volume changes to be calculated but not timed sample volume changes.

Another issue with the system that had an influence on results was that the thermometer measuring the condensed overhead outlet temperature from the heat exchanger did not accurately depict the water temperature. The condensed overhead did flow past this thermometer since it was collected in the collection tanks; however, the temperature always read between 75-80oF with no changes.

The average values of the calculated heat transfer coefficients were deemed to be higher than the typical overall heat transfer coefficients found in literature. This deviation can be attributed to uncertainties within the process. The most notable uncertainty is the potential condensation of the vapor boiled off in the 1st effect before it reached the 2nd effect. This could have altered the steam rate in the 2nd effect and produced a different overall heat transfer coefficient for the system. Also, heat loss was deemed negligible through the 2 effects because the transfer lines were insulated. This assumption may not be completely accurate, and heat loss may have occurred in the transfer pipes.

6. Conclusions and Recommendations

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The suitability of replacing the current double-effect evaporator used for concentrating high fructose corn syrup by an old evaporator found in storage was investigated due to a scheduled shutdown for routine maintenance of the current double-effect evaporator. To simplify the process and to understand the effects of the different variables, the study was performed using water instead of high fructose corn syrup as the non-concentrated solution and concentrated product. In order for the study to be performed, the height of the sight glass level against volume was calibrated as shown in Figures 4 and 5. The operation of the double-effect evaporator was characterized for a wide range of steam and vacuum pressures, and the values can be found in Table 5. It was found that the steam pressure and 2nd effect pressure have a positive correlation with the 2nd effect condensed vapor flow rate, with the steam pressure being the most influential parameter. In addition, steam efficiency is greatly affected by changes in steam and vacuum pressures. Namely, an increase in steam pressure and vacuum pressure increases the steam efficiency. Moreover, the two parameters have the same significance in steam efficiency. The heat transfer coefficients in the evaporator heat exchangers were determined for different vacuum and steam pressures as tabulated in Table 5. The patterns indicate an increase in heat transfer coefficient with increasing steam pressure for both effects. On the other hand, increasing the vacuum pressure causes a decrease in the heat transfer coefficient and a decrease in the vacuum pressure causes an increase in heat transfer coefficient for both effects. Furthermore, based on data collected the heat transfer coefficient for the heat exchanger increases with increasing steam and vacuum pressures.

6.1. Recommendations

The results outlined in this report were derived from a few data points due to time constraints; therefore, it is recommended that more experiments be performed with repetitions so that conclusions could be drawn more reliably and accurately. For this same reason, a medium level experiment should be performed to accurately characterize the operation of the double-effect evaporator. It is also recommended that the double-effect evaporator be operated at a maximum steam pressure (30-35 in Hg) as it provides the greatest heat and reduces the time frame required for evaporation to start. Based on the data collected, it was concluded that a vacuum pressure of 8 in Hg be used for experimental purposes as it provides steady state conditions. Based on literature, it was concluded that a feed-forward double-effect evaporator would operate most effectively and efficiently for this process. As this is the current mode of operation for the investigated double-effect evaporator, no changes need to be made in its setup.

For the U.S. Coast Guard’s scaled up evaporator, it is recommended that the chemical Careclean F Descalex P, produced by Marine Care, be used to effectively clean the evaporator. This product, containing sulphamic acid, is a heavy duty acid descaler that is both safe and effective to use by the personnel in the ship and provides fast cleaning without damaging other parts of the evaporator. It is recommended that the powder version be used since it does not require an extra circulation pump and the powder can just be added to the water steam feed to the evaporator.

6.2 Evaporator Setup

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When considering the optimum feed system, the advantages and disadvantages of the options must be considered. Forward-feed effect systems are generally associated with higher operating costs since the cold feed must be heated to the highest operating temperature. Conversely, this setup generally produces the most concentrated value product. However, this setup can lead to more significant scaling issues due to the increased viscosity of the product, producing a poor quality heat transfer coefficient.

The backward-feed arrangement has a lower operating cost, but a higher equipment cost due to the multiple pumps required to work against the pressure drop of the system. This feed arrangement is also useful for viscous solutions, but does not produce the product concentration that is achievable in a forward-feed system.5

Parallel-feed splits the feed stream into each effect, and removes concentrated product from each effect. The economy of this method of operation is between that of forward- and backward-feed effects, but the concentrated product is not up to standards of the forward-feed, as it produces more of slurry.5

Finally, a single-effect evaporator uses more steam upfront to produce a concentrated product similar to that of a multiple-effect system. Generally, this steam cost is not an efficient use of money and multiple-effects are used for large-scale production.2

Based on the relative advantages and disadvantages of each, a multiple-effect system is the most economical for large-scale production. The feed system depends on the required concentration of the product. In this case, the goal is to concentrate a 10% solution to a minimum of 90%. It was thus concluded that this degree of concentration is significant enough to warrant the system producing the most concentrated liquor. This system has the potential for scaling issues, but these risks are deemed to be allowable to meet the concentration requirements. Because this unit will be run on a temporary basis, it is not economically responsible to invest the higher startup cost into a backward-feed system to achieve lower operating cost for a short period of time. Thus, forward-feed is the optimum decision for this case.

6.3 Scaling

As was requested per an email from the U.S. Coast Guard, an investigation was carried out to determine possible solutions to remove heavy scale built up on an evaporator of one of the Coast Guard’s ships. This ship is located in the Caribbean and relies on its evaporator to produce potable water from the sea water. Currently, the evaporator has scaled up in an area that can only be accessed by cutting a hole in the unit with a torch, and the Coast Guard only has a citric acid cleaner available. A cleaner such as this is only useful for mild cleaning situations and not capable of removing heavy scale. After researching several potential chemicals that could be used to remove the scale, the decision was made to use a chemical produced by Marine Care called Careclean F Descalex P.

When boiling sea water in an evaporator, calcium carbonate and magnesium hydroxide are the two most prevalent compounds produced that can collect on the evaporator surface and lead to fouling.11 Calcium sulfate can also be produced but typically in a smaller quantity. Both calcium carbonate and magnesium hydroxide scale are alkaline and can be treated with acidic compounds. Sulfuric acid is efficient at removing scale of these types; however, handling

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sulfuric acid requires extreme care in handling and storage and should only be used by operators with proper training.12 Sulfuric acid would be an ideal solution to the Coast Guard’s situation, but may not be the ideal solution for safety purposes.

A more practical solution to this scenario was found in Careclean F Descalex P, a product made by Marine Care out of Holland.13 This product is a heavy duty acid descaler that contains sulphamic acid. The descaler comes as either a light-brown colored liquid or as a completely water-miscible powder with the powder method being the preferred variety for this situation; the liquid version would require an extra circulation pump for the product while the powder can just be added in the water feed stream to the evaporator. The acid powder is highly effective even at room temperature. However it provides quickest scale removal results when added as close to its maximum temperature of 60oC as possible. This cleaner can be used with most common metals (except for zinc, aluminum, and galvanized materials) and is designed to not harm clean, non-scaled surfaces of the metal.13 Treatment length depends on the thickness and type of scaling but usually take less than 24 hours. An additional treatment with Careclean Alkaline Extra product may also be used to remove any calcium sulfate that may remain in the evaporator.14

Other solutions were found for preventing scaling from occurring, but these are not ideally used for removing built up scale. Once the ship is successful in removing the scale in their evaporator, further investigation should be taken to find the most cost effective solution for maintaining the evaporator to prevent scale buildup. Examples of such scale prevention compounds include polyphosphate-lignosulfonate mixtures, sodium polymethacrylate, and surfactant compounds such as N-lauryliminodiacetic acid and any of its ammonium or alkali metal salts.12 Each of these has its own advantages for different situations. Sodium polymethacrylate would be a sufficient solution for routine scale prevention in this particular case since it is works best in temperatures up to 240oF and against calcium carbonate.12,15

Another ideal solution would be to use a treatment program such as one designed by Nalco. Nalco’s Nalfleet Maxi-vap Plus treatment service was specifically designed for seawater evaporators to prevent scale formation.16 This would prevent the Coast Guard from running into this kind of situation again as it helps maintain the evaporator over time.

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7. Nomenclature

Symbol DefinitionA1 1st Effect Heat Transfer Area (ft²)A2 2nd Effect Heat Transfer Area (ft2)Aex Condenser Heat Transfer Area (ft2)Cp Specific Heat CapacityF Feed MassHv Vapor Specific Heathl Liquid Specific HeatL Liquid Mass flow ratemcw Mass flow rate of cold waterq1 Heat transfer rate of effect 1q2 Heat transfer rate of effect 2qex Heat transfer rate of heat exchangerU1 Effect 1 overall heat transfer coefficientU2 Effect 2 overall heat transfer coefficientUex Heat Exchanger overall heat transfer coefficientΔT1 Temperature change in Effect 1ΔT2 Temperature change in Effect 2ΔTcw Temperature change in cold water streamΔTlm Log Mean Temperature Differenceλ Latent Heat of Vaporization

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8. References

1. Miletic, Marina. "Double Effect Evaporator Replacement - Rotation 1 and 2." Letter to Chris Brockman, Sudipto Guha, Cody Jensen, Victor Shum, Shannyn Stuart, Plant Supervisors. 23 Aug. 2009. MS. Roger Adams Chemical, INC., Urbana, Illinois.

2. Geankoplis, Christie J. Transport processes and separation process principles (includes unit operations). Upper Saddle River, NJ: Prentice Hall Professional Technical Reference, 2003. Print.

3. Earl, R.L. "Multiple Effect Evaporation." Unit Operations in Food Processing. New Zealand Institute of Food Science and Technology. Web. 1 Sept. 2009. <http://www.nzifst.org.nz/unitoperations/evaporation2.htm>.

4. McCabe, Warren L. Unit operations of chemical engineering. Boston: McGraw-Hill, 2005. Print.

5. Price, R. M. "Evaporation." Lecture. Www.cbu.edu. Christian Brothers University. Web. 1 Sept. 2009. <http://www.cbu.edu/~rprice/lectures/evap1.html>. Print.

6. Miletic, Marina. Double Effect Evaporator Manual. 2009. MS. University of Illinois Urbana-Champaign, Champaign.

7. Takada, Msaharu. Double-Effect Evaporator. Sasakura Engineering Company, Limited, assignee. Patent 787163. 20 Mar. 1979. Print.

8. Miletic, Marina. CHBE 430 Coursepack. 2009. MS. University of Illinois Urbana-Champaign, Champaign.

9. Mack, B., Medrano, M., Pulley, J. Double-Effect Evaporator Replacement Investigation Workplan. 2009. University of Illinois Urbana-Champaign, Champaign.

10. "Heat Exchanger Analysis." Uprm.edu. Universidad de Puetro Rico. Web. 30 Sept. 2009. <www.me.uprm.edu/o_meza/.../Heat%20Exchanger%20Analysis-3.doc>.

11. Langelier, W. F., D. H. Caldwell, W. B. Lawrence, and C. H. Spaulding. "Scale Control in Sea Water Distillation Equipment - Contact Stabilization." Industrial & Engineering Chemistry 42.1 (1950): 126-30. ACS Publications. American Chemical Society. Web. 28 Sept. 2009. <http://pubs.acs.org/doi/abs/10.1021/ie50481a034>.

12. Block, Jacob, and Nelson S. Marans. Inhibition of scale on saline water heat exchange surfaces with iminodiacetic acid compounds. W.R. Grace & Co, assignee. Patent 3,981,779. 21 Sept. 1976. Print.

13. "Careclean F Descalex P." Marine Care B.V. Marine Care Baltic. Web. 28 Sept. 2009. <http://www.marinecare.nl/pic/carecleanfdescalexp.pdf>.

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14. "Careclean Alkaline Extra." Marine Care B.V. Marine Care Baltic. Web. 28 Sept. 2009. <http://www.marinecare.nl/pic/carecleanalkalineextra.pdf>.

15. Evaporator Saline Feed Water Treatment for Scale Control. W.R. Grace & Co, assignee. Patent 3,444,054. 13 May 1969. Print.

16. "Marine transportation chemicals: water treatment for evaporators/boilers/diesel engine cooling systems." Nalco Company - Water Treatment and Process Chemical Technologies. Nalco Company. Web. 28 Sept. 2009. <http://www.nalco.com/asp/industries_served/marine/marine.asp>.

17. "Physical characteristics of water." Thermexcel. June 2003. Web. 28 Sept. 2009. <http://www.thermexcel.com/english/tables/eau_atm.htm>.

18. "Properties of Saturated Steam - Imperial Units." Engineering ToolBox. Web. 28 Sept. 2009. <http://www.engineeringtoolbox.com/saturated-steam-properties-d_273.html>.

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9. Appendix

Appendix A – Calculation Method 21

Appendix B – Experimental Raw Data 22Table B.1. Experimental data recorded for each trial 22Table B.2. Calculated data for each piece of equipment 23

Appendix C – Regression Charts 24Figure C.1. Vacuum Pressure vs. U1 24 Figure C.2. Steam Pressure vs. U1 24Figure C.3. Vacuum Pressure vs. U2 25Figure C.4. Steam Pressure vs. U2 25Figure C.5. Vacuum Pressure vs. Uhe 26Figure C.6. Steam Pressure vs. Uhe 26Figure C.7. Vacuum Pressure vs. Efficiency 27Figure C.8. Steam Pressure vs. Efficiency 27Figure C.9. Vacuum Pressure vs. Condensate Flow Rate 28Figure C.10. Steam Pressure vs. Condensate Flow Rate 28Table C.1. Steam and Vacuum Pressure vs. U1 Regression Data 29Table C.2. Steam and Vacuum Pressure vs. U2 Regression Data 30Table C.3. Steam and Vacuum Pressure vs. Uhe Regression Data 31Table C.4. Steam and Vacuum Pressure vs. Efficiency Regression Data 32Table C.5. Steam and Vacuum Pressure vs. Condensate Flow Rate Regression Data 33

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Appendix A – Calculation Method

The following equations are used to calculate the flow rates, heat transfer, and heat transfer coefficients for the two effects and the heat exchanger in the system:

1.) Evaluation of the condensing heat exchanger:2

a. Cold Water Stream: qex=mcw*Cp*ΔTcw (5)

b. Vapor-Liquid Stream: qex= m2out*Cp*ΔTvl+ΔHv (6)

c. Heat Exchanger: qex = Uex*A*ΔTlm (7)

Where qex is the heat transfer in the heat exchanger, A is the heat transfer area, m cw is the mass flow rate of the cold water, Cp is the heat capacity of the stream, T is the temperature change of the streams passing through the heat exchanger, and Tlm is the log mean temperature difference. Uex represents the heat transfer coefficient, and Hv and hl represent enthalpy constants.

2.) Evaluation of Effect 2:2

a. S2*2=q2=V2Hv+L2hl-Fhf (solve for S* and q2) (8)

b. q2=U2*A2*ΔT2 (solve for U2) (9)

c. S2=q2/2 (Solve for S2) (10)

Similar to the heat exchanger previously calculated, q2 represents the heat transfer coefficient, A represents the heat transfer area, S2 is the mass flow rate of the steam, T is the temperature difference, 2 is the latent heat of the steam, and L2, V2, and F are mass flow rates of the vapor effect. U2 represents the overall heat transfer coefficient of the effect. Hv, hl, and hf are enthalpy values for the system.

3.) Evaluation of Effect 1:2

a. S1*1=q1=V1Hv+L1hl (solve for S* and q1) (11)

b. Q1=U1*A1*ΔT1 (solve for U1) (12)

c. S1=q1/1 (Solve for S1) (13)

The calculations for the 1st effect are similar to the 2nd effect, but values for the 1st effect are used instead. They represent similar quantities just are centered around different systems.

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Appendix B – Experimental Raw Data

Table B.1. Experimental data measured for each trialRun # 1 2 3 4 5 6Vacuum Pressure (in Hg) 0 10 17 17 17 8Steam Pressure (psig) 10 10.8 29.5 31.1 24.5 30Steam Temp (°F) 254 254 255 252 250 250Feed Temp (°F) 62 62 77 63 63 63Effect 1 Temp (°F) 79.4 73.6 77 82.3 79 78Effect 1 Steam Rate (ml/min) - - 356.25 325 277 343.3Effect 1 Vapor Temp (°F) 200 201 217 215 212 223Effect 2 Temp (°F) 74.3 75.6 76.9 75.5 78.6 80.1

Effect 2 Steam Rate (ml/min) - - 214.17 200 177 220Effect 2 Initial Volume (gal) 7.600 7.150 7.950 6.000 6.365 7.150Effect 2 Final Volume (gal) 7.600 7.150 6.630 4.960 4.720 6.600Effect 2 Vapor Temp (°F) 82.3 126 165.4 164.5 165 173.8Effect 2 Pressure (in Hg) 0 -10 -17 -15 -15 -8Cold Water Flow Rate (gpm) 1.76 1 1.6 2.4 2.4 2.4Cold Water Inlet Temp (°F) 63 63 64 63 63 63Cold Water Outlet Temp (°F) 63 63 140 120 115 85Collection Tank Volume (gal) 0.000 0.000 1.110 1.030 1.639 0.480Collection Tank Temp (°F) 80 82 85 85 85 85Condensate Flow Rate (ml/min) - - 265 274 160 200

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Table B.2. Calculated data for each piece of equipment17,18

Run #   1 2 3 4 5 6Effect 1 ∆T (°F) 54 53 38 37 38 27  Area (ft^2) 1.8 1.8 1.8 1.8 1.8 1.8  Hv (btu/lb) 1150 1150 1150 1150 1150 1150  hl (btu/lb) 180 180 180 180 180 180  hf (Btu/lb) 47.43 41.62 45.04 50.31 47.02 46.03  λ (btu/lb) 942.5 942.5 942.0 944.0 945.5 945.5  V1 (lb/min) - - 0.47 0.44 0.39 0.48  L1 (lb/min) - - 1.09 0.94 0.71 0.87  S1 (lb/min) - - 0.78 0.71 0.61 0.75  q1 (Btu/min) - - 736.82 673.61 575.04 712.67

 U (Btu/min °F ft^2) - - 10.77 10.11 8.41 14.66

Effect 2 ∆T (°F) 117.7 75.0 51.6 50.5 47.0 49.2  Area (ft^2) 3.3 3.3 3.3 3.3 3.3 3.3

  Hv (btu/lb)1150.

01146.

4 1143.6 1143.6 1143.6 1147.2  hl (btu/lb) 180.0 170.5 163.6 163.6 163.6 172.5  hf (Btu/lb) 42.36 43.62 44.92 43.53 46.61 48.12  λ (btu/lb) - - 998.61 999.16 998.86 993.46  V2 (lb/min) - - 0.58 0.60 0.35 0.44  L2 (lb/min) - - - - - -  S2 (lb/min) - - 0.47 0.44 0.39 0.48  q2 (Btu/min) - - 469.58 438.75 388.18 479.87

 U (Btu/min °F ft^2) - - 2.76 2.63 2.50 2.96

  Steam efficiency - - 74.39% 84.31% 57.76%58.26

%Heat Exchanger ∆Tcw (°F) 0 0 76 57 52 22  ∆Ta (°F) 19.3 63.0 25.4 44.5 50.0 88.8  ∆Tb (°F) 10 10 86 67 62 32

 Log mean temp. dif. 14.14 28.80 49.69 54.98 55.79 55.65

  Area(ft^2) 16 16 16 16 16 16

 Heat Cap.(Btu/lb°F) 1 1 1 1 1 1

  λ (btu/lb) 970.0 975.9 980.0 980.0 980.0 974.7

 U (Btu/min °F ft^2) 0.00 0.00 1.27 1.29 1.16 0.49

  q(Btu/min) 0.00 0.00 1010.9 1137.3 1037.5 438.9823

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8 6 9

Appendix C – Regression Charts

Vacuum Pressure Line Fit Plot

6.0

7.0

8.0

9.0

10.0

11.0

12.0

13.0

14.0

15.0

16.0

0 5 10 15 20

Vacuum Pressure (in Hg)

U1

(B

tu/m

in*f

t^2

*F)

U1

Predicted U1

Figure C.1. Vacuum Pressure vs. U1

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Steam Pressure Line Fit Plot

8.0

9.0

10.0

11.0

12.0

13.0

14.0

15.0

0 10 20 30 40

Steam Pressure (psig)

U1 (

Btu

/min

*ft^

2*F

)

U1Predicted U1

Figure C.2. Steam Pressure vs. U1

Vacuum Pressure Line Fit Plot

2.4

2.6

2.8

3.0

0 5 10 15 20

Vacuum Pressure (in Hg)

U2 (

Btu

/min

*ft^

2*F

)

U2Predicted U2

Figure C.3. Vacuum Pressure vs. U2

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Steam Pressure Line Fit Plot

2.4

2.6

2.8

3.0

0 10 20 30 40

Steam Pressure (psig)

U2 (

Btu

/min

*ft^

2*F

)

U2Predicted U2

Figure C.4. Steam Pressure vs. U2

Vacuum Pressure Line Fit Plot

0.0

0.2

0.4

0.6

0.8

1.0

1.2

1.4

0 5 10 15 20

Vacuum Pressure (in Hg)

Uh

e (

Btu

/min

*ft^

2*F

)

Uhe

Predicted Uhe

Figure C.5. Vacuum Pressure vs. Uhe

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Steam Pressure Line Fit Plot

0.0

0.2

0.4

0.6

0.8

1.0

1.2

1.4

0 5 10 15 20 25 30 35Steam Pressure (psig)

Uh

e (

Btu

/min

*ft^

2*F

)

Uhe

Predicted Uhe

Figure C.6. Steam Pressure vs. Uhe

Vacuum Pressure Line Fit Plot

50%

55%

60%

65%

70%

75%

80%

85%

90%

0 5 10 15 20

Vacuum Pressure (in Hg)

Effi

cie

ncy

Efficiency

Predicted Efficiency

Figure C.7. Vacuum Pressure vs. Efficiency

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Steam Pressure Line Fit Plot

50%

55%

60%

65%

70%

75%

80%

85%

90%

0 10 20 30 40Steam Pressure (psig)

Effi

cie

ncy

Efficiency

Predicted Efficiency

Figure C.8. Steam Pressure vs. Efficiency

Vacuum Pressure Line Fit Plot

0

50

100

150

200

250

300

0 5 10 15 20

Vacuum Pressure (in Hg)

Condensate

Flo

w R

ate

(m

l/m

in)

Condensate Flow Rate

Predicted Condensate Flow Rate

Figure C.9. Vacuum Pressure vs. Condensate Flow Rate

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Steam Pressure Line Fit Plot

0

50

100

150

200

250

300

0 5 10 15 20 25 30 35

Steam Pressure (psig)

Condensate

Flo

w

Rate

(m

l/m

in)

Condensate Flow Rate

Predicted Condensate Flow Rate

Figure C.10. Vacuum Pressure vs. Condensate Flow Rate

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Table C.1. Steam and Vacuum Pressure vs. U1 Regression Data

SUMMARY OUTPUT

Regression StatisticsMultiple R 0.9831R Square 0.9665Adjusted R Square 0.8994Standard Error 0.8388Observations 4

ANOVAdf SS MS F Significance F

Regression 2 20.28 10.14 14.41 0.18Residual 1 0.70 0.70Total 3 20.98

Coefficients Standard Error t Stat P-value Lower 95% Upper 95% Lower 95.0% Upper 95.0%Intercept 9.27 5.66 1.64 0.35 -62.67 81.21 -62.67 81.21Steam P 0.31 0.17 1.80 0.32 -1.88 2.50 -1.88 2.50Vacuum P -0.49 0.11 -4.36 0.14 -1.91 0.94 -1.91 0.94

RESIDUAL OUTPUT

Observation Predicted U1 Residuals1 10.12 0.662 10.61 -0.503 8.57 -0.164 14.66 0.00

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Table C.2. Steam and Vacuum Pressure vs. U2 Regression Data

SUMMARY OUTPUT

Regression StatisticsMultiple R 0.9304R Square 0.8656Adjusted R Square 0.5967Standard Error 0.1224Observations 4

ANOVAdf SS MS F Significance F

Regression 2 0.10 0.05 3.22 0.37Residual 1 0.01 0.01Total 3 0.11

CoefficientsStandard Error t Stat P-value Lower 95% Upper 95% Lower 95.0% Upper 95.0%Intercept 2.39 0.83 2.89 0.21 -8.11 12.89 -8.11 12.89Steam P 0.03 0.03 1.08 0.48 -0.29 0.35 -0.29 0.35Vacuum P -0.03 0.02 -1.90 0.31 -0.24 0.18 -0.24 0.18

RESIDUAL OUTPUT

Observation Predicted U2 Residuals1 2.66 0.102 2.71 -0.073 2.53 -0.024 2.96 0.00

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Table C.3. Steam and Vacuum Pressure vs. Uhe Regression Data

SUMMARY OUTPUT

Regression StatisticsMultiple R 0.9999R Square 0.9998Adjusted R Square 0.9995Standard Error 0.0082Observations 4

ANOVAdf SS MS F Significance F

Regression 2 0.43 0.22 3227.42 0.01Residual 1 0.00 0.00Total 3 0.43

CoefficientsStandard Error t Stat P-value Lower 95% Upper 95% Lower 95.0% Upper 95.0%Intercept -0.81 0.06 -14.69 0.04 -1.51 -0.11 -1.51 -0.11Steam P 0.02 0.00 12.07 0.05 0.00 0.04 0.00 0.04Vacuum P 0.09 0.00 79.64 0.01 0.07 0.10 0.07 0.10

RESIDUAL OUTPUT

ObservationPredicted Uhe Residuals1 1.27 0.012 1.30 0.003 1.16 0.004 0.49 0.00

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Table C.4. Steam and Vacuum Pressure vs. Efficiency Regression Data

SUMMARY OUTPUT

Regression StatisticsMultiple R 0.9926R Square 0.9852Adjusted R Square 0.9557Standard Error 0.0273Observations 4

ANOVAdf SS MS F Significance F

Regression 2 0.05 0.02 33.39 0.12Residual 1 0.00 0.00Total 3 0.05

CoefficientsStandard Error t Stat P-value Lower 95% Upper 95% Lower 95.0% Upper 95.0%Intercept -0.75 0.18 -4.09 0.15 -3.09 1.59 -3.09 1.59Steam P 0.04 0.01 6.88 0.09 -0.03 0.11 -0.03 0.11Vacuum P 0.02 0.00 6.15 0.10 -0.02 0.07 -0.02 0.07

RESIDUAL OUTPUT

ObservationPredicted EfficiencyResiduals1 0.77 -0.022 0.83 0.023 0.57 0.014 0.58 0.00

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Table C.5. Steam and Vacuum Pressure vs. Condensate Flow Rate Regression Data

SUMMARY OUTPUT

Regression StatisticsMultiple R 0.9879R Square 0.9760Adjusted R Square 0.9279Standard Error 14.5851Observations 4

ANOVAdf SS MS F Significance F

Regression 2 8638.03 4319.01 20.30 0.16Residual 1 212.72 212.72Total 3 8850.75

Coefficients Standard Error t Stat P-value Lower 95%Intercept -400.61 98.45 -4.07 0.15 -1651.48Steam P 18.16 3.00 6.06 0.10 -19.90Vacuum P 6.96 1.95 3.57 0.17 -17.80

RESIDUAL OUTPUT

Observation Predicted Condensate Flow Rate Residuals1 253.59 11.412 282.65 -8.653 162.77 -2.774 200.00 0.00

34