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Thermodynamic Release Scenario Modeling and Air Dispersion Modeling for Incident Prevention, Mitigation and Emergency

Response Planning at an Acrolein Storage and Transfer Facility

R. Maqbool Qadir, PE Enpro Solutions, Inc.

6500 Dublin Blvd., Suite 215, Dublin, CA 94568 Qadir@enprosolutions.com

James A. Westbrook

BlueScape Environmental 16870 West Bernardo Drive, Suite 400, San Diego CA 92127

jwestbrook@bluescapeinc.com

© R. Maqbool Qadir and James A. Westbrook

Prepared for Presentation at American Institute of Chemical Engineers

2014 Spring Meeting 10th Global Congress on Process Safety

New Orleans, LA March 30 – April 2, 2014

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UNPUBLISHED

AIChE shall not be responsible for statements or opinions contained in papers or printed in its publications

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Thermodynamic Release Scenario Modeling and Air Dispersion Modeling for Incident Prevention, Mitigation and

Emergency Response Planning at an Acrolein Storage and Transfer Facility

R. Maqbool Qadir, PE Enpro Solutions, Inc.

6500 Dublin Blvd., Suite 215, Dublin, CA 94568 Qadir@enprosolutions.com

James A. Westbrook

BlueScape Environmental 16870 West Bernardo Drive, Suite 400, San Diego CA 92127

jwestbrook@bluescapeinc.com

Keywords: acrolein, loss of containment, LOC, pressure relief valve, PRV, thermal expansion, hydrostatic pressure, worst case scenario, alternate release scenario

Abstract Thermodynamic release modeling and vapor dispersion modeling are performed to quantitatively evaluate the hazards posed by loss of containment scenarios identified in hazard analyses for an acrolein storage facility. Offsite consequence modeling requires that the amount, duration and state (vapor, liquid, dense gas) of the release must be estimated. Typically, thermodynamic models are not expected to explain how a release occurs. However, in this case by analyzing the temperature-pressure-time behavior of the acrolein liquid and vapor space in response to a fire or thermal runaway, thermodynamic modeling provides insight into the actual sequence of events that must take place to cause a release. In this paper we look at the effect of non-condensable nitrogen in the headspace and the impact of liquid thermal expansion when evaluating LOC hazards. Liquid thermal expansion could cause a potentially catastrophic breach of the container and turns out to be the dominant mechanism for causing LOC rather than a rise in acrolein vapor pressure. Thus it is important to evaluate thermal expansion hazards and the relief system design should take the need for hydrostatic relief into account. The amount of acrolein released as a result of relieving hydrostatic pressure due to thermal expansion is estimated to be substantially higher than would be expected from a vapor release. 1. Introduction Acrolein is a highly toxic, flammable and reactive chemical widely used in agriculture and oil field chemical formulations. This case study presents the results of thermodynamic modeling and offsite consequence analysis (OCA), including worst case

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scenario (WCS) and alternate release scenario (ARS), conducted for an acrolein storage facility in New Iberia, Louisiana. Hazard analysis and OCA are required by federal, state and local regulations for facilities storing greater than threshold amounts of regulated toxic or flammable substances. Applicable federal regulations include the Risk Management Program (RMP) under the Clean Air Act [1] and the United States Occupational and Health Safety Administration’s (OSHA) Process Safety Management (PSM) Standard [2]. The hazard analysis is useful for facility siting, accident prevention and emergency response planning even when stored quantities of hazardous materials are below regulatory thresholds for a particular facility. In this case study, the results of the modeling were used for two separate but related purposes:

1. To make engineering and planning decisions about onsite emergency response and incident management.

2. To generate estimates of potential release rates and expected durations of acrolein emissions for use as input for air dispersion modeling to conduct the OCA.

The site (Figures 1: Site Vicinity Map) was owned and operated by a chemical distributor to the oil and gas industry and was used for the storage of full and partially empty acrolein containers, ranging in size from rail car containers to smaller day tanks. The acrolein containers included a nitrogen blanket in each tank to prevent oxidation of the contents and were protected by rupture disk and pressure relief valve (PRV) combinations.

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2. Hazard Identification and Assessment Hazard analysis was conducted in accordance with regulatory guidelines and available technical resources. These included United States Environmental Protection Agency (US EPA) and OSHA guidance, acrolein material safety data sheets (MSDS), the safety literature concerning acrolein, container and storage tank construction details, company operating procedures, and hazard and operability studies (HAZOP). This analysis identified several scenarios likely to result in loss of containment (LOC) i.e. accidental acrolein releases, including the following WCS and ARS: 1. WCS - rupture of a single container to a contained liquid pool. 2. Hose rupture during transfer ‐ liquid spill. 3. Vent line rupture during:

a) container filling - vapor phase release. b) tank purging ‐ vapor phase release.

4. External fire and pressure relief valve operation – liquid, vapor or 2-phase release. 5. Rapid polymerization and pressure relief valve operation ‐ liquid, vapor or 2-phase release.

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The HAZOP identified the LOC scenarios on a qualitative basis. The nature of the release(s) that might occur in response to an external fire or hazardous polymerization event in Scenarios 4 and 5 in particular were poorly understood. A number of questions needed to be answered to evaluate the hazards posed by the scenarios. For instance offsite consequence modeling requires that the amount, duration and state (vapor, liquid, dense gas) of the release must be estimated. Figure 2 illustrates the acrolein container, hazard scenarios 4 and 5, and the questions to be answered. It was clear that an appropriate level of system modeling was required to gain a quantitative understanding of how the system would behave.

Figure 2: Acrolein Containers, Hazard Scenarios 3. Consequence Modeling: Thermodynamic analysis of thermal expansion due to fire or hazardous polymerization 3.1 System Definition and Modeling Approach

The container system under study is shown in Figure 2. Initially, the vessel is 95% full and nitrogen blanketed to 1 bar pressure. The system PRV has a set pressure of 110 psig and the rupture disk burst pressure is 125 psig. For calculations, data for a 1-inch, Kunkle 6010 series pressure relief valve are used. The PRV has a typical blow-down value (the pressure at which it will reseat) of 6% of the set pressure. This means that the relief valve would continue relieving until the pressure in the vessel was reduced to 6% below the

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PRV set pressure, i.e. to 103.4 psig. Clearly a storage tank exposed to fire or subject to internal heating due to hazardous polymerization would tend to heat up and become over-pressured. But what phenomena are involved and what is the sequence of events that lead to a burst disk and PRV operation? The first law of thermodynamics for a non-flow situation is:

∆ [Eq. 1] Q = heat absorbed by the system W= work performed by the system ∆ E = Energy change of the system In the absence of external work this reduces to:

∆ [Eq. 2]

Thus all the heat absorbed by the tank and its contents would result in increasing the internal energy of the system. As the liquid in the tank is heated up, three mechanisms can be identified that would cause the pressure to increase inside the acrolein container: liquid expansion upon heating would reduce the headspace and compress the nitrogen, acrolein vapor pressure would increase, and the expanding liquid could ultimately lead to liquid filling the entire container and exerting hydrostatic pressure on the vessel walls. The modeling approach is shown in Figure 3 and thermodynamic models and calculations for each of these processes are described below.

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Figure 3: Modeling Approach 3.2 Fire Scenario Heat Source, API Standard 521, API, RP 520 method An external pool fire model developed and used by the United States Department of Transportation (USDOT) in a tank car study [3] is used as the heat input model. Correlations for bulk liquid heating rates and some acrolein physical properties are also taken from the DOT study. The USDOT Study uses the more conservative of two fire scenarios identified in the American Petroleum Institute (API) Standard 521 [4] i.e. one where good drainage and adequate fire-fighting equipment may not be available. The heat rate for this scenario is described by Equation 3 below. 34,500 . [Eq. 3]

Q is the radiant heat rate in btu/hr generated by a pool fire with a flame temperature of 1700 and absorbed by the wetted surface of the tank exposed to the fire. Awet is the wetted area of a horizontal cylinder with spherical ends and is given by Equation 4, taking into account the effective liquid level or holdup [5]:

180 [Eq. 4] E = effective liquid level D = diameter

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L = length B = effective liquid level angle in degrees F = environmental factor, with a value of 1 for a bare (uninsulated) tank surface Q for the specified ISO container in this scenario is calculated to be 4.45 10 btu/hr or 7.82 10 joules/min. 3.3 Liquid heat rate and time to reach specified temperature The bulk liquid heat up rate is given by the USDOT model [3]:

9.724 10 / . [Eq. 5]

where is the bulk heat-up rate in /minute.

f = a dimensionless environmental factor (1 for uninsulated tank).

= the density x heat capacity product averaged over a temperature interval of interest in J/m3 K. V = the tank volume in m3

For this scenario, is calculated to be 2.86 /min.

Time to reach a specified liquid temperature from the initial temperature [3]:

/ [Eq. 6]

3.4 Liquid Expansion and Reduction of Headspace The percent volume expansion from some initial temperature Ti to a specified bulk liquid temperature TL is given by: ∆ % 100 ∆ / [Eq. 7] ρ is the acrolein liquid density at TL , and Δρ is the density change over the temperature interval (TL - Ti). Density values for acrolein were estimated by interpolating from density data obtained at different temperatures from the DOT study [3]. The final headspace volume at any acrolein liquid temperature of interest TL is calculated as:

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∆ [Eq. 8] This reduction in volume compresses the nitrogen in the headspace raising the pressure and temperature of the gas in the headspace as calculated below. The acrolein liquid temperature that produces a total pressure in the headspace equal to the rupture disk burst pressure is calculated by iterating TL and computing the headspace pressure as described below. 3.5 Headspace Pressure and Initial Molar Composition By Dalton’s law the total pressure in the headspace is the sum of the nitrogen pressure and the acrolein vapor pressure. [Eq. 9]

Where is the vapor pressure at the temperature of liquid acrolein and is the partial pressure of nitrogen at the temperature and volume of the headspace in the vessel. At an initial temperature of 30 in both liquid and vapor phases, and pressure of 1 barg the initial mole fraction of N2 is 0.83 based on the partial pressures of the two components. At the temperature and pressure of interest the compressibility factor (ratio of the molar volume of a real gas to the molar volume of an ideal gas at the same temperature and pressure) of the mixture is about 0.95. The error in treating the gas (83% nitrogen) as an ideal gas is tolerable. We can then calculate the initial moles of N2 in the headspace before the fire using the ideal gas law.

/ [Eq. 10] The moles of acrolein can be calculated by:

/ [Eq. 11] Where the “i” subscript indicates initial conditions for the headspace. 3.6 Pressure and Temperature of N2 and acrolein in the Headspace after Polytropic Compression by Expanding Liquid The moles of N2 are unchanged and are equal to the initial state (MN2). At a temperature of TL in the liquid phase and assuming the headspace gas can be represented as an ideal gas, the pressure and temperature of nitrogen in the headspace can be calculated as follows:

2 [Eq. 12]

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2

1 [Eq. 13]

n is the polytropic constant = 1.4 for N2. Pacrolein is the vapor pressure at the temperature TL of liquid acrolein The moles of acrolein can be calculated by:

/ [Eq. 14] The total pressure of the headspace can be calculated by Dalton’s law - Equation 9. Substituting the expression for (Eq 12) in Equation 9 [Eq. 15]

The final headspace volume is given by Equation 8 as a function of the change in the liquid volume, which in turn is a function of the change in density. Density is not available as an explicit function of temperature and is estimated by interpolating between known data points. Therefore, the coupled set of Equations 7 through 13 is solved by iterating the liquid temperature TL and estimating and until Pt equals the rupture disk burst pressure Pb. The temperature calculation is then closed and the headspace volume, temperature, pressure, and molar composition of the vapor phase are fixed. Release conditions can now be estimated. Figure 4 below shows the acrolein vapor pressure and nitrogen partial pressure for a range of liquid temperature. The plot shows that the nitrogen partial pressure rises much faster than the acrolein vapor pressure as the liquid temperature increases. The PRV set pressure is exceeded at 60 .

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Figure 4: Acrolein Vapor Pressure, Headspace Nitrogen Pressure vs. Liquid Temperature 3.7 Gas Phase Release from the Headspace (constant volume process) The PRV response to excess pressure is typically of the order of seconds and the venting capacity for the particular model selected for analysis is of the order of 1300 standard cubic feet per minute (scfm). Therefore, the release period will be small compared to the heating rate of the liquid (2.86 /min), the headspace will not change appreciably, and the change in temperature and pressure of the headspace can be modeled as a constant volume process during the release. The gas/vapor phase release occurs from the rupture disk burst pressure of 125 psig to the PRV reseating pressure of 103.4 psig. No liquid flashing would occur because the total pressure in the headspace exceeds the saturation pressure of acrolein at TL. Thus the headspace pressure following the first release will be 103.4 psig. The temperature in the gas phase of the headspace is given for this constant volume process by

/ [Eq. 16] The partial pressure of the nitrogen remaining in the headspace following the release is estimated by Dalton’s law - re-arranging Equation 9 gives

[Eq. 17] The mass of N2 released = final N2 vapor space mass – initial N2 vapor space mass.

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Figure 5: Fire Scenario, thermal expansion, ISO Tank - Acrolein, Nitrogen and Total Pressure 3.8 Tank filling and Rate of Pressure Rise Generated by Thermal Expansion Tank filling occurs because of volume expansion of the acrolein as it continues to be heated by the fire. Once the tank is completely full, continued expansion generates hydrostatic pressure that can be estimated through pressure, volume, and temperature (pvt) relationships. Figure 4 and 5 show that tank full conditions are reached when the acrolein temperature reaches 70 and this occurs within 15 minutes of the start of a fire. For a single phase fluid, an exact thermodynamic relationship between molar volume, temperature and pressure is as follows [6]:

[Eq. 18]

For constant volume conditions (such as a completely full container), this may be integrated to give a relationship between the change in pressure generated by a change in temperature [6]:

∆ ∆

or ∆ /∆ [Eq. 19]

where β is the volume expansivity defined as:

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1/ [Eq. 20]

κ is the isothermal compressibility defined as:

κ 1/ [Eq. 21]

A value of κ for acrolein was obtained from literature sources, while β was estimated from volumetric data derived from known densities. The rate of pressure rise for a completely filled container was estimated by combining Equations 5 and 19 as follows: ∆

∆

∆

∆175.5 10 / . [Eq. 22]

Once the container is full, the rate of pressure rise for this scenario is calculated to be 25.40 bar/min or 368.2 psi/min. This would cause catastrophic failure in a short time unless the pressure is relieved by the PRV at a sufficient rate. 3.9 Minimum Relief Rate and Actual Relief Rate The minimum flow rate required to relieve the pressure rise can be estimated from an energy balance between the rate of heat input and removal by the liquid flowing out of the tank from the PRV. This flow rate is given by the Equation 23 where the terms are as defined in API Standard 521 [4] but the notation is different:

/ [Eq. 23]

β is the volume expansivity previously defined Q = heat rate btu/hr Cp = heat capacity btu/lbm-F S = liquid specific gravity The minimum flow rate is calculated as 5.9 gpm. The actual relief rate from the PRV is given by Equation 24 [7] omitting inapplicable or negligible correction factors for viscosity, rupture disk and back pressure.

38 . . [Eq. 24]

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Qf = flow rate in gpm A = valve effective area, inch2

PR = pressure upstream of relief valve Pa = relief valve exit or ambient pressure G = specific gravity Kd = ASME valve discharge coefficient Kp = over-pressure correction coefficient The actual relief rate at liquid full conditions is calculated at 263 gpm, far more than the minimum rate of 5.9 gpm required to prevent catastrophic failure. The temperature of the liquid released is 70 well above its normal boiling point of 53 therefore it would exceed its saturation vapor pressure and would start to flash upon release. 3.10 Vapor Generation Rate An estimate of the vapor generation rate is required for estimating the amount of vapor that would need to be relieved should the vapor pressure of acrolein exceed the PRV set pressure. Once the PRV opens an acrolein vapor release would tend to continue until fire-fighting measures reduce the temperature, and therefore the vapor pressure, to below the PRV set pressure, terminating the release. The vapor generation rate may be estimated as follows. λ = 5.18 10 joules/kg, latent heat of evaporation at the normal boiling temperature Q = 7.82 10 joules/min previously calculated heat rate The vapor generation rate in kg/minute – m': m' = Q/λ [Eq. 25] The vapor generation rate is calculated as 151 kg/min or 332 pounds/min. 3.11 Hazardous Polymerization Calculating the self-heating rate for hazardous polymerization (thermal runaway) requires experimental data and the solution of coupled differential equations representing mass and energy balances. The required data include the heat of reaction as well as chemical kinetics- reaction order and the reaction rate or reaction rate constant. Literature values for the heat of reaction were obtained but no satisfactory kinetic data could be found. The self-heating rate can also be established experimentally by direct measurement in the laboratory using accelerated rate calorimetry (ARC). We could not find literature references for ARC data either. It is not possible to estimate the self-heating rate dQ/dt (joules/min) or the rate of temperature rise dT/dt (C/min) without any experimental data. However it is possible to calculate the temperature rise corresponding to a particular conversion (fraction of material reacted) given the heat of reaction. The calculation

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results plotted below show what conversion is required for the material initially at 30 C to reach a temperature corresponding to the PRV set pressure.

Figure 6 - Acrolein Hazardous Polymerization, Liquid Temperature vs. Mass Conversion 3.12 Summary of Thermodynamic Modeling The thermal scenario occurs as a result of a fire in the facility. Initially, the vessel is 95% full and nitrogen blanketed to 1 bar pressure. The liquid in the vessel gets hot and expands, reducing the headspace volume and compressing the nitrogen. The pressure would rise until the nitrogen/acrolein vapor mixture reaches the rupture disk bursting pressure of 125 psig. The rupture disk would burst, and nitrogen with potentially some entrained acrolein vapor/liquid mixture would be released through the PRV, which has a set pressure of 110 psig. The relief valve would continue relieving until the pressure in the vessel was reduced to 6% below the PRV set pressure, i.e. to 103.4 psig. A continuing release would occur if the fire persisted and the vessel continued to heat up until the headspace pressure (nitrogen partial pressure plus acrolein vapor pressure) exceeded the PRV set pressure completely purging the nitrogen. Once the nitrogen is purged thermal expansion causes tank full conditions. The PRV relief rate is sufficient to prevent catastrophic tank failure but would result in a significant amount of liquid phase being released over a short period. The release would terminate when the external heat source was removed. The site is fully staffed and is located in the vicinity of a fire station and we would expect that a fire would be detected and reported promptly. For modeling purposes we assumed that a fire response would occur within 15 minutes time and a fire would be extinguished within 2 minutes of response.

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Release 1 (gas phase): A liquid temperature of about 60 or 140 causes thermal expansion, producing headspace pressure of 125 psig, sufficient to exceed the rupture disk burst pressure (Figures 3 and 7). The release is mainly nitrogen gas, which continues to be purged over the next 3-4 minutes to relieve built up pressure. The time required to reach this temperature is about 10.5 minutes from the start of the fire.

Release 2 (flashing liquid): Thermal expansion of the liquid as it heats up and purging of the nitrogen blanket causes the vessel to become completely full in 14-15 minutes from the start of the fire. The overpressure causes the PRV to release liquid acrolein at a rate of 263 gpm, or 1850 pounds per minute compared to a minimum relief rate necessary to prevent catastrophic failure of the tank of only 5.9 gpm. The liquid acrolein at 70 C would start to flash upon release, as its normal boiling point is 53 C. For modeling purposes, it is assumed that the total release lasts 2 minutes after fire response occurs until the heat source is removed and further releases are terminated. The total amount released would be 526 gallons, or 3700 pounds of acrolein.

Figure 7: Thermodynamic Modeling Summary

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4. Air Dispersion Modeling 4.1Air Modeling Approach An OCA was conducted for the WCS and the selected ARS. A worst-case release scenario is defined as the release of the largest quantity of a regulated substance that results in the greatest distance to a toxic endpoint. Under the US EPA RMP regulations, the toxic endpoint for acrolein is 0.0011 mg/L or 0.5 ppm. The endpoint is the Emergency Response Planning Guideline-2 (ERPG-2) value, which assumes 60-minute exposure. The worst-case release scenario does not consider the likelihood that such an event will occur. Thus, the assumptions used for the worst-case scenario are highly conservative. The alternative release scenario is defined as a release that is more likely to occur than the worst-case release scenario and that reaches an endpoint offsite, unless no such scenario exists. The OCA completed for the acrolein storage facility followed US EPA guidance documents: RMP Offsite Consequence Analysis Guidance [8], and RMP Guidance for Chemical Distributors [9]. The US EPA’s RMP*CompTM [10] and Areal Locations of Hazardous Atmospheres (ALOHA) [11] software programs were used to complete chemical source-term calculations and estimate the furthest distance to the acrolein toxic endpoint. 4.2 Worst-Case Release Scenario The worst-case release scenario for the acrolein containers would be catastrophic breach of a single ISO shipping container. Although highly unlikely, for the worst-case analysis, the US EPA requires an assumption that the entire contents of the container are released over a 10-minute period. In the case of a catastrophic container breach, the release of liquid to the containment area is assumed to occur instantaneously. To estimate the furthest distance to the acrolein toxic endpoint, the RMP*CompTM model was used. The worst-case release scenario modeling methodology and results are described below. 4.3 Dispersion Modeling Methodology In order to complete the worst-case dispersion modeling with RMP*CompTM, it was assumed that all 4,790 gallons (34,783 pounds) of acrolein is released from one acrolein ISO container into the diked storage area. This diked area was considered as a passive mitigation measure for the worst-case analysis. The acrolein release rate was calculated by assuming the entire contents of the tank are instantaneously released to a pool in the diked area. From this pool, acrolein will evaporate into the air over a 95-minute period. It was also assumed that 100% of the released liquid is acrolein. The pool size of a 4,790 gallon spill was calculated in order to determine the acrolein release rate to be used in the worst-case modeling scenario. Using the formula for volume of an oblique pyramid (Volume = (1/3) * base area * height), the pool size (base area) for a 4,790 gallon spill was found to be about 3,986 ft2. The method was used

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because the diked area is sloping and the diked area would exceed the spill area. This information was entered into the RMP*CompTM model. As directed by US EPA OCA guidance [8], the maximum recorded temperature in New Iberia (from the National Climatic Data Center) of 103F was used for acrolein storage and pool temperature in the modeling analysis. Given this is an extreme temperature recorded on a single day, the acrolein storage contents are very unlikely to be at this temperature level. An ambient air temperature of 77F, a minimum wind speed of 1.5 m/s, and “F” stability was assumed by RMP*CompTM in this analysis. The topography of the region is generally flat and unobstructed, so a rural terrain setting was used. 4.4 Distance to the Acrolein Endpoint Using RMP*CompTM, the maximum distance to the acrolein endpoint was calculated to be greater than 25 miles (40 km). This is the greatest distance that the RMP*CompTM model will report. In cases where the maximum distance to endpoint is greater than 25 miles, the user is instructed to report the distance as “25 miles” in the RMP. As required by US EPA, the worst-case release scenario assumed worst-case meteorological conditions described above. RMP*CompTM model input assumptions and output for the worst-case release scenario are shown in Table 1 below. Table 1 ACROLEIN STORAGE WORST-CASE RELEASE SCENARIO Model Input Assumptions and Results Model Used RMP*CompTM Substance Released 100% Acrolein Scenario Considered Single ISO container breach, liquid release Modeled Definition Ground-level finite duration Passive Mitigation 3986 ft2 spill area, 6 inch dike; fully contained Release Duration Instantaneous Release Rate 479.7 lbs/min Release Height 0 m Ambient Air Temperature 77F (model default) Relative Humidity 50% (model default) Wind Speed 1.5 m/sec (model default) Stability Class F (model default) Acrolein Storage Temperature 103F (NCDC Climate Summary 1948-2000) Surface Terrain Rural Acrolein Toxic Endpoint 0.0011 mg/L (ERPG-2 basis) Release Rate to Outside Air 367 lb/min evaporation rate Evaporation Time 94.7 minutes Distance to Toxic Endpoint >25 miles (>40 km)

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4.5 Offsite Receptor Endpoints The estimated residential population within the worst-case circle is about 29,268 people, as compiled from 2000 U.S. census data. The sensitive and/or public receptors that are found inside the impact circle consist of private residences, schools, and hospitals/health care facilities. The environmental receptors within the worst-case impact circle consist of federal land. 4.6 Alternative Release Scenario The US EPA RMP guidance maintains that an alternative release scenario must be included with the RMP analyses if the facility has a worst-case impact on the surrounding public. The purpose of the alternative scenario is to provide information on a more reasonable release scenario of the chemical using more typical meteorology conditions. This alternative analysis therefore provides more reasonable information on potential impacts due to a chemical release of acrolein as compared to the worst-case study. Three release scenarios were reviewed to select one alternative scenario: 1) rupture of a transfer hose with spill of acrolein liquid from the one-inch hose into a 1 centimeter deep pool, 2) rupture of the one inch vent line from a ISO container to the scrubber, during a) field tank container filling, b) purging, with release of acrolein vapor, and 3) rapid polymerization within one ISO container due to external heating and release of acrolein vapor. Acrolein emission calculations were completed and modeling performed using the ALOHA dispersion model for each of the scenarios. The results are listed in Table 2 below. The release durations shown do not account for active mitigation that could be employed to contain the spill. Table 2: Alternative Release ALOHA Modeling Results

Scenario Release Form Release Rate

Maximum Release Duration

Distance to Acrolein Endpoint

1. Transfer hose rupture, field tank filling

Liquid into pool, evaporation

0.70 lb/min

Up to 20 minutes, depending upon response

0.29 miles

2. Vent line rupture, line purging

Vapor 0.48 lb/min

Up to 5 minutes, depending upon response

0.24 miles

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3. Rapid polymerization, field tank

Vapor 87.3 lb/min

28 seconds 2.2 miles

While a transfer line or vent line rupture would be a much more likely scenario, the polymerization release scenario would be more difficult to mitigate or control, and would lead to 8-10 times greater distance to acrolein toxic endpoint than would the other scenarios. Therefore, the alternative release scenario selected for the facility was a vapor release due to a polymerization in a field tank. In either instance, heating would cause rapid pressurization in the container that could lead to PRV disk rupture. An instantaneous release of acrolein vapor in the container headspace could be followed by periodic releases of acrolein vapor, until the ISO container involved is cooled through a fire or heat suppression system. The selected alternative release scenario, modeling methodology, and results are described below. The ALOHA model was used to complete the alternative scenarios modeling. 4.7 Dispersion Modeling Methodology In order to complete dispersion modeling with ALOHA, the acrolein release rate was assumed to be a nearly instantaneous release of ACR vapor during PRV disk rupture. The release amount was estimated to be 87.3 lbs over less than a one minute release period, the minimum release period considered in ALOHA. Additional periodic releases were not modeled. It is assumed that any additional release amounts would be about the same or less than the initial release. As required by the USEPA in the Risk Management Program Guidance for OCA [8], an average ambient air temperature of 66.9F, an average wind speed of 4.0 m/s, and “D” stability was assumed by ALOHA in this analysis. An average relative humidity of 74% was used. The topography of the region is generally flat and unobstructed, so a rural terrain setting was used.

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Figure 8: ALOHA Output, Selected Alternate Release Scenario 4.8 Distance to the Acrolein Endpoint Using ALOHA, the maximum distance to the acrolein endpoint was calculated to be about 2.2 miles (3,541 meters). As required by US EPA, the alternative release scenario assumed average meteorological conditions described above. Source term calculations and ALOHA model input assumptions and output for the alternative release scenario are shown in Table 3 below. Table 3 ACROLEIN STORAGE ALTERNATIVE RELEASE SCENARIO Model Input Assumptions and Results Model Used ALOHA Substance Released 100% Acrolein Scenario Considered PRV failure from rapid heating, vapor release Modeled Definition Instantaneous

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Passive Mitigation PRV valve at 125 psig Active Mitigation Fire or temperature suppression Release Duration 28 seconds (modeled as 1 minute) Release Rate 87.3 lb/minute, 1.45 lb/second Release Height 0 m Ambient Temperature 66.9F (NCDC Climate Summary 1948-2000) Relative Humidity 74% (local meteorological data) Wind Speed 4.0 m/sec (local meteorological data) Stability Class D (assumed per OCA guidance) Surface Roughness Rural Acrolein Toxic Endpoint 0.0011 mg/L (ERPG-2 basis) Distance to Acrolein Endpoint 2.2 miles (3,541 meters) 4.9 Offsite Receptors The RMP rules require offsite receptors within the circular area formed by the distance to the acrolein endpoint to be identified, including residential populations, and public and environmental receptors. Generally, the public receptors located within the 2.2 mile radius are the private residences and businesses to the west of the facility.

5. Conclusions Typically, thermodynamic models are not expected to explain how a release occurs. However, in this case by analyzing the temperature-pressure-time behavior of the acrolein liquid and vapor space in response to a fire or thermal runaway, thermodynamic modeling provides insight into the actual sequence of events that must take place to cause a release i.e. rupture disk burst at 60 , followed by PRV venting nitrogen, followed by tank full conditions at 70 resulting in a liquid-phase release. Specific insights gained from modeling included the following:

1. The presence of the non-condensable nitrogen in the headspace caused the pressure to exceed the rupture disk burst pressure much faster than would have been expected on the basis of acrolein vapor pressure alone. This implies that it is important to consider the effect of non-condensables in the system when evaluating LOC hazards.

2. In this case liquid thermal expansion is the dominant mechanism for causing LOC and could cause a potentially catastrophic breach of the container. Thus it is important to evaluate thermal expansion hazards and the relief system design should take the need for hydrostatic relief into account.

3. The amount of acrolein released as a result of relieving pressure due to thermal expansion (1850 pounds/min) is estimated to be substantially higher than would be expected from a vapor release (332 pounds/min). Consideration should be given to providing a safe relief discharge location such as a sump to minimize the spread of hot liquid acrolein in case of hydrostatic relief.

4. The feasibility of reducing the maximum fill level for acrolein containers should be explored to minimize the potential for tank full conditions in case of fire.

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5. Modeling results imply that, in case of fire or hazardous polymerization the acrolein container rupture disk could burst in a little over 10 minutes and emergency response should take place in 15 minutes or less to prevent serious releases.

6. Appropriately located air monitoring systems could give early warning of a release and enable an automated response.

7. An appropriately located water deluge system could greatly reduce the potential for a fire to cause a release and would tend to mitigate the release if one does occur.

Facility process control and emergency response planning decisions made using release scenario modeling and OCA simulation results included the following: Identifying setback distance to maintain separating acrolein from flammable

products. Sizing concrete berm and adding sumps to contain spills and reduce the size of the

liquid pool thus minimizing the amount of acrolein released to air. Covered sumps could potentially provide safe relief system discharge locations.

Implementation of mitigation measures, including a water deluge system for fire suppression and reducing evaporation.

Perimeter air monitoring sensors and alarms. 6. References [1] USEPA. “Chemical Accident Prevention Provisions”, 40 CFR, 68, Subparts C

and D. U.S. Environmental Protection Agency

[2] OSHA. “Standard for Hazardous Materials - Process Safety Management of Highly Hazardous Chemicals”, 29 CFR, 1910.119, U.S. Occupational Safety and Health Administration

[3] USDOT., “Railroad Tank Car Relief Valve Requirements for Liquid PIH Lading”, Department of Transportation (2003), DOT/FRA/ORD-03/21, September 2003

[4] API. “Guide for Pressure-relieving and Depressuring Systems: Petroleum petrochemical and natural gas industries”, American Petroleum Institute (2007), API Standard 521, Fifth Edition, January 2007

[5] Crosby. “Pressure Relief Valve Engineering Handbook”, Crosby Valve, Inc., Technical Document No. TP-V300, May 1997

[6] Smith, J.M., and Van Ness, H.C., Introduction to Chemical Engineering Thermodynamics, Third Edition, 1975, McGraw Hill

[7] API. “Sizing, Selection, and Installation of Pressure-Relieving Devices in Refineries, Part I - Sizing and Selection”, American Petroleum Institute (2000), API Standard 520, Part I, Seventh Edition, January 2000

GCPS 2014 __________________________________________________________________________

[8] USEPA. “Risk Management Program Guidance for Offsite Consequence Analysis”, March 2009, EPA 550-B-99-009, http://www.epa.gov/oem/docs/chem/oca-chps.pdf

[9] USEPA. “Risk Management Program Guidance for Chemical Distributors”, March 2009, EPA 550-B-00-005, http://www.epa.gov/osweroe1/docs/chem/oca-chps.pdf

[10] USEPA. RMP*Comp.www.epa.gov/osweroe1/content/rmp/rmp_comp.htm.

[11] http://response.restoration.noaa.gov/sites/default/files/ALOHA_Tech_Doc.pdf , Areal Locations of Hazardous Atmospheres (ALOHA) Version 5.4.4, August 2013, USEPA and NOAA