Ventilation requirements for chiller rooms. (2023)

INTRODUCTION

Section 8.11.5 of ASHRAE Standard 15 (ASHRAE 2010) states:

"The mechanical ventilation required to relieve an accumulation of refrigerant due to leakage or rupture of the system shall be capable of removing air from the machinery space in not less than the following quantity:

Q = 100 [square root of G]

(Q = 70 [square root of G])

where Q = the air flow in cubic feet per minute (liters per second) and G = the mass of refrigerant in pounds (kilograms) in the largest system, a portion of which is located in the engine room.'

The Standard 15 User's Manual (Fenton and Richards 2002) states that the table of ventilation requirements on which the formula is based was probably developed before 1930. In fact, Brown (2005) shows that this table first appeared in the Explosive and Hazardous Trades of the New York Municipal Code, 1927 (Section 220 of Article 18) and also gives some other antecedents in the development of ventilation requirements. This table did not appear in the ASRE Handbook until 1939, when it was accepted with modifications by the then B9 Committee. The modifications consisted of higher ventilation rates for systems with more than 1000 lbs of refrigerant (Brown 2005).

The Standard 15 equation has several fundamental deficiencies: 1) the equation does not account for variations in maximum acceptable refrigerant concentration (or recommended concentration limit - RCL); 2) The equation does not take into account differences in refrigerant properties such as boiling point, vapor pressure and molecular weight; and 3) it does not account for different room sizes.

This research program was undertaken in an effort to improve the Standard 15 equation. First, accident history related to engine room refrigerant leaks was reviewed to help determine conservative but reasonable scenarios for future accidents. Next, a differential equation was developed to estimate the ventilation rate required to limit the coolant concentration to the maximum limits for the proposed accident scenario. Finally, a simplified equation was derived for easy application and implementation in the standard.

The aim was to provide a technical rationale for setting ventilation requirements in chiller rooms to maintain safety and to develop equations or other methods to calculate required ventilation rates. To achieve this goal, the ventilation strategies and rates required to minimize health hazards from a leak are based on a realistic accident scenario, reflect the principles of dilution, take into account the recommended concentration limits and physical properties of each refrigerant, and are simple enough to be used in standards and building codes to determine ventilation requirements for refrigeration machinery spaces.

BACKGROUND

A multi-database overview shows release scenarios, including quantity, duration, location in the system and cause of the release. Several have offered their opinion on realistic release rates to calculate ventilation needs.

Reported Accidents

The search for historical refrigerant leaks provided the direction to define realistic and appropriate accident scenarios on which ventilation requirements can be based. All recorded accidents in North America and Europe have been accounted for with accidents related to engine room refrigerant leaks, including the extent of the leak, the cause and any associated personal injury or property damage. Four main databases were searched:

* MARS (Major Accident Reporting System) (MAHB 2001): Database on "major accidents" reported under Seveso, OECD and UN-ECE. Managed by the Major Accidents Bureau (MAHB). The current MARS database contains over 700 accidents and near misses collected from European Union Member States since 1982. Between May 1991 and December 2001, 17 accidents were investigated.

* NRC (The National Response Center) (NRC 2010): NRC provides annual data files containing data on incidents (oil, chemical, radiological, biological and etiological spills throughout the United States and its territories). Table 1 shows the number of reports to the NRC (2010) regarding the refrigerants suspected above as likely to be in engine rooms. Reportable Quantity (RQ) (US EPA 2001) is the threshold quantity at which 24-hour cumulative refrigerant loss must be reported to the NRC by law. Ammonia is the only refrigerant with a low enough reportable quantity to be meaningful. Therefore, almost all accidents are ammonia. Some of the accidental releases reported by the NRC are also included in the ARIP database.

Table 1. Reported Solid Source Releases (NRC) from 1 January 2004 to 31 December 2009Refrigerants Number of Incidents Reportable Quantity (lb)R-11 1 5000R-12 2 5000R-22 13 *(1)R-123 0 *(1 )R-134a 2 Not ListedR-245f 0 Not ListedR-407a 0 Not ListedR-410a 0 Not ListedR-417a 0 Not ListedR-422a 0 Not ListedR-290 (Propane) 0 10,000R-601 0 Not ListedR-601a 0 unlistedR-717 (Ammonia) 6010 100R-744 146 unlisted* (1) Chemical which is reportable under Section 313 and Section 6607 of the Pollution Prevention Act.

* ARIP (The Accidental Release Information Program) (US EPA 2010): ARIP is a database compiled by the EPA that focuses on accidental releases at fixed installations that resulted in off-site consequences or environmental damage. All non-routine spills of oil and chemicals must be reported (to NRC, Coast Guard or EPA regional offices). The US Environmental Protection Agency (EPA) compiles these reports in the Emergency Response Notification System (ERNS) database. Significant accidents are selected from the ENS Accidental Release Information Program (ARIP) database and a questionnaire is sent to the affected facility consisting of 23 questions about the facility, the circumstances and causes of the incident, and the practices and technologies used to address the incident Accidental Release Prevention that existed prior to the event and were added or modified as a result of the event. The ARIP database was searched for indoor refrigerant release accidents for the period 1990-1999. Because the EPA obtains information from the NRC database, most of the accidents found in the ARIP database also appear in the NRC database. However, the ARIP database provides more detailed information about the cause of the release, as well as the amount and duration of the release. Therefore, where possible, the information extracted from the ARIP database has been supplemented with the NRC data to provide as much information as possible for each incident. If the NRC database conflicts with the ARIP database, the ARIP database is assumed to be more accurate. 223 accidents were recorded that met the criteria.

* OSHA (Occupational Safety and Health Administration) (OSHA2010): The OSHA database focuses on accidents affecting workers and contains accident investigation summaries. The Accident InvestigationSummary database provides details of accidents affecting workers and contains limited information about the cause of a chemical spill, the duration of the spill, or the amount of the spill. A search was conducted and indoor refrigerant release accidents between January 1990 and February 2010 were extracted, yielding 94 applicable accidents. Eighty of the reported releases were of ammonia. Only limited information is available on the amounts released.

* NIOSH (National Institute for Occupational Safety and Health) (NIOSH 2010): The Centers for Disease Control (CDC) and NIOSH report fatalities in some accidents (mostly workplace) under the Fatality Assessment and Control Evaluation (FACE) program. Only one death could be found related to a cold leak. In 1992 in Alaska, an assistant ice rink manager died of asphyxiation while trying to stop a CFC-22 leak in a compressor room.

All databases listed have been analyzed to identify accident scenarios that are conservative but have a reasonable probability of occurring in practice. The ARIP/NRC databases are the most comprehensive. For most reported accidents, the release quantity, release duration and release parameters (gas, liquid, etc.) are listed. Unfortunately, mainly ammonia accidents are recorded (see Table 2.1 for an explanation). . The most common release amounts are between 100 and 10,000 pounds. The released amounts were classified and counted. The probability distribution of the emission rate is shown in Figure 1.

[FIGURE 1 OMIT]

Table 2. Probability distribution of emission rate from ARIP/NRC database (ammonia) Probability of 1% 2% 5% 10% Exceeding total emission 2.99 kg/s 2.29 kg/s 1.13 kg/s 0.61 kg /s Rate (Metric) Total Emission 395.9 302.4 149.7 80.4 lb/minRate (English) lb/min lb/min lb/min

The databases show that accidents were more common before 1993, as shown in Figure 2. In 1992, OSHA issued the Process Safety Management (PSM) of Highly Hazardous Chemicals Standard (OSHA 1992), which provides requirements for the management of hazards associated with highly hazardous chemical processes. For refrigeration systems, the standard applies to systems containing 10,000 pounds of ammonia. Any facility with the threshold amount of ammonia must have a PPP program in place. All new plants must develop and implement a plan before importing ammonia in excess of the threshold.

[FIGURE 2 OMIT]

Table 2 lists the total emission rate for specific exceedance probabilities. These emission rates were obtained by linear interpolation. One percent of all accidents assessed had an overall emission rate greater than 2.99 kg/s (395.9 lb/min). Five percent of the accidents had a total emissions rate greater than 1.13 kg/s (149.7 lb/min).

So far, no distinction has been made according to the type of release (gas, liquid or both). Figure 3 shows the proportion of different release types from the ARIP/NRC database. Gaseous ammonia was released in 75% of the reported accidents and only 13% of the accidents involved a pure liquid release. However, the boiling point of ammonia is 240 K (-28 °F). A build-up of liquid ammonia would create vapor in most engine room environments due to heat transfer from the bottom to the liquid. Therefore, even a purely liquid release would have a gaseous component. On the other hand, a liquid release could be confused with a gaseous release because the sudden pressure drop at the release site causes part of the liquid jet to vaporize and the remaining liquid can vaporize almost immediately.

gas 75% liquid 13% gas and liquid 12%

Figure 4 shows the percentages for different causes of the reported releases. Equipment failure is the leading cause of accidents. Calculating the average emission rate for all release causes was 0.27 kg/s (35.7 lb/min) for operator error, 0.27 kg/s (35.7 lb/min) for equipment failure and 0.16 kg/s ( 21.2lb/min) for others. Therefore, no distinction can be made between the magnitudes of releases due to device failure or due to operator error.

Operator error 27% Device error 68% 0 more 5% Figure 4 Reason for release from ARIP/NRC database. Note: Table from pie chart.

The ARIP database also gives the place of publication. The MARS and OSHA databases were also evaluated. Both the OSHA and MARS databases have large percentages of unknown locations. Although no emission rate information is given, release locations are listed. Figure 5 shows the percentage of occurrence of different release sites. The most common leak point is at a valve. The second most likely location is the pipeline.

(a)A 46%B 23%C 15%D 4%E 12%(b)A 23%B 23%C 18%D 6%E 6%F 24%(c)A 27%B 27%C 6 %D 1%E 7%F 32% Figure 5 Release location from (a) ARIP/NRC, (b) MARS and (c) OSHA databases. A: valves; B: Piping, C: Compressor/Process Vessel, D: Pump, E: Other, and F: Unknown. Note: Table from pie chart.

Although the total mass of the refrigerant is not as important as the emission rate, it does play a role in the analysis as it represents a significant limit to the amount of refrigerant that can be released. This leads to an emission time where increased ventilation can slow the increase in refrigerant concentration in the room, but until the leak stops or the entire mass is released, the ventilation will not clean the room.

The upper limit is generally defined by the catastrophic failure of the largest ship in the engine room. The amount of refrigerant stored in this tank could be fully released if a large crack occurs at the bottom. The released refrigerant would consist of the initial flash gas followed by the remaining liquid. The pool of liquid on the floor of the engine room would generate vapor due to heat transfer to the liquid from the floor. After some time, as the soil cools and approaches the equilibrium temperature at atmospheric pressure for the coolant, the rate of vapor production would decrease depending on heat transfer from the air above the liquid surface. Another failure that could cause a large release would be a large break in a fluid supply line. If the flow is not stopped, a very large volume of liquid could be expelled through the line. Several incidents occurred in this way, leading to the release of large quantities of ammonia.

In commercial buildings with chiller rooms with water chillers, the smaller release events are similar to those seen with ammonia in industrial plants. These leaks are caused by valve packing failures, mechanical seal leaks, tube and pipe joint failures, corrosion failures, or improper operating or maintenance procedures. Refrigerant storage tanks are not typically used in conjunction with chillers, so the amount of refrigerant in these systems (R-11, R-22, R-123, R-134a, etc.) is significantly lower. Consequently, the largest release of refrigerant from water chillers is significantly smaller than the potentially large release from an industrial facility.

Refrigerant Release Types

Generally. Brown (2005), Stoecker (1998), Richards (1986), and others suggest that a failure consisting of a rupture of a 1/2-inch high-pressure, high-temperature fluid line is the most likely design-worst-case scenario is. Depending on the refrigeration system, such as piping and valves, the leak rate can be calculated by applying flow principles in conjunction with the thermodynamic properties of the refrigerant.

Seidl and Taylor (2005) study a leak from a 0.25 inch (6.35 mm) hole. Using R-22 and temperatures and pressures of 40°F (4.4°C) and 83 psia (572 kPa) and 100°F (37.8°C) and 210 psia (1448 kPa), the leak rate was calculated to be 3.5 lb/min (0.026 kg/s) and 8.5 lb/min (0.064 kg/s) respectively. For a half inhole at the higher temperature and pressure, the leak rate was calculated to be 34 lb/min (0.257 kg/s). They recommend a leak rate of 0.11 kg/s (15 lb/min), arguing that similar non-toxic refrigerants have similar leak rates and that a 0.25 inch (6.35 mm) hole is not an unreasonable estimate of a hole or accidental drilling is.

Refrigerant releases are generally jet streams that are all vapor, all liquid, or a mixture of vapor and liquid (i.e., two-phase), the designer determines the aeration rate. The four variables that affect vapor leak rate are: pressure, temperature, state of the refrigerant (vapour or liquid), and physical size of the hole or opening. The variables that typically cause higher leak rates are high pressure, high temperature, large orifices, and/or fluid release. Low pressure, low temperature, small orifices, and/or vapor release typically typify lower leak rates. This is not entirely the case in all situations, but it correctly indicates general trends. For example, a high pressure, high temperature liquid leaking through a large orifice would be large. In contrast, a leak of low-pressure, low-temperature vapor through a small orifice would be relatively small. Fluids under high pressure will provide a greater leak rate due to the high density and greater flow rate through the orifice. High temperature liquids also increase the leak rate due to the higher percentage of flash vapor generated. From another perspective, coolant leaks are larger, and therefore more difficult to manage, than vapor leaks through the same orifice, pressure, and temperature. High pressure, high temperature, liquid leaks therefore generate the largest amount of vapor (volume) and thus pose the greatest safety risk in engine rooms.

Fluid leaks are possible at the following temperature and pressure combinations: high pressure and high temperature, high pressure and low temperature, and low pressure and low temperature. At high pressure and high temperature, the flash steam accounts for about 20% to 40% of the total mass flow released, depending on the actual pressure and temperature. The flashing of the vapor cools the remaining liquid to its thermodynamic equilibrium temperature at atmospheric pressure. The temperature of the coolant feeding the leak also drops in proportion to its pressure. When the upstream pressure drops to atmospheric pressure, the leak rate through the orifice is nearly zero. However, depending on the location of the leak, the refrigerant may still be present in the source container or pipe. The spilled liquid absorbs heat from surfaces it comes in contact with and consequently evaporates while cooling that surface. When enough liquid has leaked, it can accumulate on the floor of the engine room and only produce a small amount of vapor after generating large amounts of vapor that cool the floor. At this point, the heat that vaporizes the refrigerant liquid comes from only a small amount of heat energy still remaining in the ground and the surrounding air.

When the liquid is at a low temperature, the amount of flash vapor is significantly less (of the order of 10% by mass) and hence the vapor generated is less. But when the pressure is high, the spilled fluid is quickly replenished and the spilled fluid creates vapor that cools the surfaces it comes in contact with in the engine room. A low pressure liquid will also produce flash vapor at a similarly low fraction, but will not be vigorously replenished like a high pressure liquid.

The amount of vapor that will exit an orifice depends only on the orifice size and the upstream pressure, assuming the flow is not restricted. The upstream temperature affects the leak rate just as it affects the density of the upstream vapor. However, if the upstream pressure is greater than the critical pressure (and depends on the hole size), then the flow will be choked and not dependent on the upstream pressure. Under choked flow conditions, upstream temperature and pressure changes affect flow rate in the same way they affect upstream vapor density.

Liquid jet releases. If the hole causing the release is at or below the liquid-vapor interface, liquid will escape. The pressure in the container or pipe resulting from the state of thermodynamic equilibrium forces the liquid through the hole. With refrigerants, the sudden drop in pressure causes part of the liquid jet to turn into vapor. It is therefore a two-phase flow. Because refrigerants generally experience some degree of liquid-to-vapour flashing, releases of pure liquids are not considered except to predict the liquid flow approaching the hole or rupture.

The flow of liquid from a container, e.g. a storage tank, through a hole depends on the pressure difference between the inside and the outside of the container and on the head that the liquid develops over the hole (AIChE 1996):

[E.sub.l] = [c.sub.o] [A.sub.h] [[rho]l][[2((p-[p.sub.o])/[[rho ]]) + 2 g [H 1]] 1/2] (1)

The shape of the vessel also affects the flow rate of a liquid through a hole. For a vertically aligned vessel, the liquid height can be calculated according to (AIChE 1996).

[Hl] = 4[Vl]/[pi][d2] (2)

If the vessel is horizontal, then (AIChE 1996),

[H.sub.l] = d/2 (1 - cos[θ]l]) (3)

[V.sub.1] = L[d.sup.2]/4 (&thgr;) – sin(2[&thgr;].sub..) / 2) (4)

Reservoir tanks are typically located outside of refrigeration machine rooms. Therefore, a jet of liquid exiting a pipe is a more appropriate scenario. Fluid flow out of a pipe depends on the difference between fluid pressure and ambient pressure, as well as the pressure drop through fittings (e.g. elbows, tees and valves) and wall friction (ASHRAE 2009). Therefore,

[E.sub.l] = [A.sub.h][[rho]l][[2((p-[p.sub.o])/([[rho].sub.l] ([ff]L/D + [[SIGMA].sup.K])))]1/2] (5)

The friction loss depends on the Reynolds number Re of the flow and on the roughness height H of the pipe wall surface (ASHRAE 2009):

[ff] = 8[[[[(8/Re)].sup.12] + 1/[(A+B).sup.1,5]].sup.1/12] (6)

[A.sub.h] = [[2.457.ln(1/([(7/Re)].sup.0.9] + (0.27[epsilon]/D)))].sup.16] (7)

B = [(37530/Re)16] (8)

Re = 4[E.sub.l] / [[rho].sub.l][pi]Dv (9)

Since the friction loss f depends on the mass emission rate [Q l ], the mass emission rate has to be calculated iteratively. For high Reynolds numbers, i.e. fully turbulent flow, the friction factor becomes independent of the Reynolds number.

Two-phase jet releases. The release of liquid refrigerant, which involves a significant pressure drop, causes some of the liquid to vaporize into vapor. Assuming that the expansion of the liquid is adiabatic, that air does not mix with the expanding liquid and that the expansion process is reversible, the amount of flashing liquid can be calculated using an isentropic energy balance (API 1996):

[Ef]/[El] = [Sl1] – [Sl3]/[DELTA][Swap](10)

If one neglects the kinetic energy of the liquid expansion, one can calculate the quantity of the flashing liquid with the isenthalpic balance. The heat of vaporization is obtained entirely from the enthalpy of the released liquid. It can be shown that the fraction of liquid flashed is (API 1996):

[Ef]/[El] = [cp](T – [Tb])/[Hvap] (11)

The release of liquid refrigerant from high pressures involves a significant change in kinetic energy, so the isenthalpic compensation approach may be unsuitable for calculating the amount of liquid splashed. Neglecting the kinetic energy of liquid expansion causes an overprediction of flashing liquid. According to the Manual for Modeling Hypothetical Accidental Releases to the Atmosphere (API, 1996), Equation 11 should be used in cases where a saturated liquid is released. With refrigerant mixtures, the expansion process is much more complex and the calculations must include information about the vapor-liquid equilibrium conditions of the mixture.

Two-phase flow releases from vessels and pipes with choked flow remain an area of ​​research. Although empirical relationships have been proposed, they only apply to simple situations: refrigerants made up of individual components. One approach that has been used for single-component two-phase flows is the homogeneous equilibrium flow model (HEM), which is based on the assumptions:

* Homogeneous liquid-vapour mixture

* Thermal equilibrium between liquid and vapor phase

* No slippage between liquid and vapor phases

* Isentropic expansion process

The HEM has been applied in determining the release of liquids at choked flow rates, but releases of refrigerant have not been found in the literature. Several researchers have used this approach to predict laboratory results, but only (Sallet 1990) has predicted the flow rates of several refrigerants including: R-290 (propane), ammonia, carbon dioxide, and R-22.

The monograph (AIChE 1996) reports another approach to predicting two-phase choked flows using analytical expressions. The thermodynamic conditions must be appropriate to allow an overall fluid flow equation similar to Bernoulli's equation.

evaporation basin. A two-phase release allows liquid to pool on the bottom while releasing a supercooled vapor at the same time. The emission rate for the two mechanisms adds up. An equation considered for liquid spills is given below (US EPA 1992):

Q = 6,94 × [10 –7 ] (1 + 0,0043 [[[T 2 ] – 273,15] 2 ]) × [U 0,75 × [As] × [MW] x [Vp] / [Vph] (12)

[Vph] = exp(76,858 – 7245,2/[T2] – 8,22ln([T2]) + 0,0061557 [T2] (13)

The vapor pressure is calculated using the Clausius Clapeyron equation:

[Vp] = 10,1325 * exp ([Hvap][Mw]/R (1/[Tb] – 1/[T2])) (14)

The liquid basin surface area and superficial velocity can all be adjusted to obtain specific evaporation rates for alternative emission scenarios and room sizes.

The evaporation of a pool of liquid is a time-varying process. Initially the pool is very small; therefore the evaporation rate is insignificant. However, as time progresses, the pool of liquid becomes larger, which increases the rate of evaporation. Eventually the pool is large enough that equilibrium is reached and the evaporation rate approaches the liquid release rate (total mass emission rate minus emission rate of ejected liquid). The limiting size of the pool is the floor area of ​​​​the room or the area of ​​\u200b\u200ba dike, if used.

Vapor Jet clearances. The release of vapor jets can be well approximated assuming isentropic expansion (reversible and adiabatic). The estimates provided give reasonable temperatures of the initial vapor plume and velocity of the vapor entering the plume. For vapor releases that result in a large change in velocity, the isentropic approximation has been shown to be preferable to the isenthalpic approximation (Moranand Shapiro, 2000).

A jet of steam forms upon the development of a small hole that creates a path for the coolant to flow through and enter the surrounding air in the engine room. When the pressure in the vessel or tube containing the refrigerant vapor exceeds the critical pressure, the flow rate through the hole is sonic (equal to the speed of sound) and the flow is defined as "choked". No matter how much greater the refrigerant pressure is above the critical pressure, the flow rate remains constant. Only the density in the container or pipe (throughput temperature and pressure) influences the flow rate. Assuming that the refrigerant vapor behaves like an ideal gas, the critical pressure ratio is

p / [po] [greater than or equal to] [((k+1)/2)k/(k-1)](15)

Values ​​for k generally range from 1.2 to about 1.5, and the pressure ratio defined by Equation 16 is about 2 for most gases and vapors.

The choke flow condition for an ideal gas or vapor based on isentropic flow is given by

E = [co][Ah][[[[rho]k(2/(k+1))](k+1)/(k-1)]].sup .1/2](16)

The discharge coefficient values ​​vary from 0.6 to almost one and are due to non-ideal flow influences. Note that when the pressure in a vessel is above the critical pressure, the temperature drops somewhat, causing the density to decrease, which in turn reduces the flow rate.

When the pressures are less than the critical pressure ratio, the choked flow condition no longer persists and the mass flow rate is dependent on the pressure inside the tube vessel. The flow rate is given by

E = [c0][Ah][[2p[rho](k/(k-1))([([p0]/p)2/k ]-[([p o ]/p) (k+1) /k])] 1/2] (17)

Note that the mass flow is lower than with choked flow.

For a given release, the rate of pressure change within the container or tube depends on the specific situation governing the release. Either an adiabatic or an isothermal flow process may be suitable. An example of isothermal flow would occur when there is a small leak in a pipe, as the cooling caused by the expansion is counteracted by the frictional heating and heat transport through the pipe from the outside. This would keep the temperature in the tube constant; however, at relatively large vapor release flow rates, the expansion process cools the internal volume of the tube or vessel and the adiabatic process is the appropriate choice. In addition, if the pipe or vessel is insulated, then again the adiabatic process is the appropriate choice. In the situations where the isothermal process persists, the internal temperature is constant. With the adiabatic flow method, the pressure drops when the refrigerant is discharged. The pressure can be estimated using conveniently small time increments and calculating the mass released over each time step. For each subsequent time step, subtract the mass that has escaped from the mass in the vessel or pipe and recalculate the pressure. In this way, even the case of isothermal flow and heat transfer to or from the vessel or pipe can be accounted for using reasonable values ​​for the convection heat transfer coefficient.

Further elaborations on calculation methods for steam releases are contained in the monograph (AIChE 1996), in which methods for releases from pipe bursts developed by Wilson (1981) are discussed for the first time. The developed concept involves defining the initial vapor mass in the tube and treating the release as an isothermal vapor release characterized by an expression consisting of the sum of two exponential terms. The exponential terms are actually time constants that depend on the physical size of the pipe and the size of the fracture.

PUBLISH SCENARIOS

The purpose of this section is to present the method used to develop a realistic vapor release rate for an engine room based on refrigerant type and historical information on accidental releases.

There are five factors that affect vapor release rates as previously discussed: release orifice size; refrigerant condition (gas or liquid); properties of the refrigerant; pressure and temperature. Different combinations of these factors produce different vapor release rates. Below are the combinations ordered from high to low vapor release potential as stated by Brown (2005):

* Release Scenario 1: Liquid, High Pressure, High Temperature

* Release scenario 2: gas, high pressure, high temperature

* Release scenario 3: liquid, high pressure, low temperature

* Release Scenario 4: Liquid, Low Pressure, Low Temperature

* Release Scenario 5: Gas, low pressure, low temperature

Of most interest is the realistic worst case scenario and the provision of background information to use in selecting a realistic release rate to be used for ventilation room design. Figure 6 shows a schematic of an atypical refrigeration cycle that identifies where these release scenarios might occur in the process.

[FIGURE 6 OMIT]

Release Scenario 1: Liquid, High Pressure, High Temperature

A high pressure, high temperature liquid release is a two-phase flow. The pressure inside the container or tube resulting from the state of thermodynamic equilibrium forces the liquid through the hole. The sudden drop in pressure at the exit point causes part of the liquid jet to turn into vapor. The remainder of the liquid collects on the engine room floor at boiling point temperature (Brown, 2005) and then slowly warms to ambient temperature. The vapor emission rate is therefore calculated in 3 steps: 1) the total liquid mass emission rate from the hole is determined; 2) the proportion of flashing liquid is calculated; and 3) the evaporation rate from the surface of the liquid pool is calculated.

Several assumptions must be made regarding release size, release area, release temperature, and air velocity over the coolant pool. The liquid refrigerant pool size was assumed to be the equilibrium pool size (the evaporation rate was limited to no more than the liquid release rate minus the flash liquid emission rate) with a liquid depth of 0.394 inches (10 mm) (US EPA 1992). As the liquid exits the tube, it expands and cools to the boiling point. For refrigerants with boiling points below ambient conditions, the liquid pool will gradually warm up, but will most likely never reach ambient temperature before fully evaporating. The liquid pool also cools the surrounding air. Therefore, the pool temperature T2 in Equation 13 was taken as the greater of 32°F (0°C) or the boiling point temperature of the specific refrigerant. This is a conservative assumption for chemicals with boiling points below 0°C (32°F) since the liquid will remain cooler than 0°C (32°F) for some time. For example calculation purposes, the air velocity over the spillway is assumed to be 50 ft/min (0.254 m/s). This would correspond to a ventilation rate of 20,000 cfm recommended by Brown (2005) for all engine rooms and a room cross-sectional area of ​​400 [ft2] (37.2 [m2]). A lower ventilation rate and smaller room size would give a similar air velocity, which is the only important input to the evaporation calculation.

To determine the total vapor emission rate, the mass flow rate of the flash liquid (using the isentropic balance) is added to the evaporation rate.

A potential valve failure point is directly downstream of a high-pressure, high-temperature liquid refrigerant storage tank, for example, a failed liquid line valve downstream of the receiver tank.

A simplified short pipe release without wall friction, valve loss coefficients and head minimizes the necessary assumptions about the cooling system. This triggering scenario is solely dependent on the diameter of the hole and the pressure difference inside and outside the pipe. Therefore, this scenario is simple yet realistic.

Equation 2 simplifies to:

[E.sub.l] = [c.sub.o][A.sub.h][[rho]sub.l][[2((p-[p.sub.])/[p.sub .l])].sup.1/2] (18)

The EPA recommends an outflow coefficient [c 0 ] = 0.6, but this coefficient depends on the flow conditions outside the orifice, where [c o ] = 1 represents flow unimpeded by the shape of the orifice (e.g. a severed pipe).

Pressure and temperature in the reservoir are kept constant during release (US EPA 1992). In reality, the loss of liquid refrigerant from the valve causes evaporation inside the tank and therefore a reduction in the temperature and pressure of the system. The resulting total liquid emission rate would therefore decrease slightly faster over time than that calculated assuming constant pressure and temperature, which is therefore the simpler and more conservative approach.

As described above, the flash emission rate of R-134a release is calculated using the isentropic method. The remaining liquid R-134a forms a pool. Pool size is time dependent and is calculated using an air velocity of 50 ft/min (0.254 m/s).

The flash liquid velocity is proportional to the total liquid release rate and slowly decreases over time. Shortly after the release, the pool size is small and therefore the evaporation rate of the pool is insignificant. However, over time the pool grows until an equilibrium is reached where the amount of liquid refrigerant supplying the pool equals the rate of evaporation, resulting in a total vapor emission rate equal to the total liquid emission rate.

The calculations described above were repeated for other hole diameters. Figures 7 and 8 show the dependency of the various calculated emission rates on the diameter of the release opening for R-134a and ammonia, respectively. Total liquid mass emission rate, vapor emission rate due to flashing, equilibrium liquid pool evaporation rate, and total vapor emission rate are plotted against hole diameter on a log-log scale for two samples: ammonia (R-717) and 1,1,1,2-tetrafluoroethane ( R-134a). The vapor released by flashing and evaporating a 0.25 inch (6.35 mm) hole accounts for about 95% (5% would exceed this figure) of the accidents found in the databases. A 0.5 inch. (12.7mm) hole exceeds 99% of accidents.

Release Scenario 2: Gas, high pressure, high temperature

This was initially rated as the second worst vapor release potential. High pressure, high temperature gas release is possible in the piping between the compressor and the condenser (see Figure 6). When the pressure in the pipe exceeds the critical pressure, the flow rate equals the speed of sound through the hole and the flow is "choked". When this is the case, the vapor emission rate is constant no matter how much greater the pressure inside the tube is above the critical pressure. The throttled state is therefore the borderline case for this scenario. This scenario has significantly lower emission rates than observed (see Table 3.2) and is not considered a reasonable case to base the room ventilation rate on.

Release Scenario 3: Liquid, High Pressure, Low Temperature

This scenario is very similar to scenario 1 described. A release of high-pressure, low-temperature liquid refrigerant is possible between the heat exchanger and the expansion valve. All the equations used in Scenario 1 can also be applied to this scenario. Insignificant differences in the resulting emission rates are expected for this scenario compared to Scenario 1, since the total liquid emission rate is driven solely by the pressure difference between the inside and outside of the container. The lower temperature of the refrigerant causes a slightly lower flash rate and a lower evaporation rate. However, once the system reaches equilibrium (i.e. the evaporation rate equals the total liquid emission rate minus the flashing rate), the temperature difference between the two scenarios has no effect and this scenario yields the same emission rates as Scenario 1.

[FIGURE 7 OMIT]

[FIGURE 8 OMIT]

Release Scenario 4: Liquid, Low Pressure, Low Temperature

A liquid release of refrigerant at low pressure and low temperature is possible between the expansion valve and the evaporator. If the pressure inside the tube or vessel falls below ambient pressure, the release will be governed by the weight of the liquid over the opening and no release can occur until the pressure equalizes. Due to the lower temperature compared to Scenario 1, the flash liquid rate is significantly lower. The size of the equilibrium basin is smaller than that for Scenario 1 because the total liquid emission rate is lower. Since the calculated emission rates are significantly lower than scenario 1, this scenario is also excluded for determining the ventilation rate.

Release scenario 5: gas, low pressure, low temperature

This scenario is very similar to scenario 2. A release of gaseous refrigerant with low pressure and low temperature is possible between the evaporator and the compressor. All the equations used in Scenario 2 can also be applied to this scenario. The resulting vapor emission rate is expected to be lower than that calculated in Scenario 2, since the total vapor emission rate is driven solely by the pressure difference between the inside and outside of the canister. A reduction in this pressure difference causes a reduction in the emission rate. When the pressure inside the pipe or vessel drops below the critical pressure, the restricted flow condition no longer persists, so the vapor flow rate depends on the pressure inside the vessel or pipe. The total mass flow rate for unchoked flows is less than that for choked flows. Since the calculated emission rates are well below scenario 1, this scenario is also excluded for determining the ventilation rate.

Recommended release scenario for ventilation room design

To account for a likely worst case scenario, it is recommended to use high temperature high pressure release. Although the two-phase release is the most realistic, accounting for all of the evaporation pond variables, including superficial velocity, pond depth, pond size (floor area), and heat transfer considerations, becomes very difficult and full of ambiguous assumptions. In addition, the emission rate for the evaporation pond can become very large. Therefore, it is believed that a drain is provided to catch any liquid that would collect. All vapor emissions are generated by the flash vapor from a 0.5 inch (12.7mm) hole. Considering only the flash fumes (no evaporation from potential liquid) this represents an accident scenario for ammonia with only a 5% exceedance probability, which seems a reasonable probability for design purposes (including evaporation this scenario would give a 1% exceedance probability). ). It is further assumed that similar probabilities would occur for other refrigerants.

VENTILATION RATE

The purpose of engine room ventilation is to dilute the room air so that the recommended concentration limit (RCL) in Tables 1 and 2 of ANSI/ASHRAE Standard 34 (ASHRAE 2010) is not exceeded. However, depending on the actual leak rate, ventilation may or may not dilute the refrigerant concentration to the RCL or lower using the current Standard 15 method. This section presents methods that a designer can use to specify an aeration rate to ensure that the concentration is below the RCL value. In the following sections, a derivation of the equations required to estimate the concentration in a machine room due to an accidental release of a refrigerant is presented.

Detector activation of ventilation system

Standard 15 (ASHRAE 2010) specifies the requirement that every chiller room contain a detector that activates an alarm and mechanical ventilation in the event of a refrigerant leak. The alarm shall activate at a refrigerant concentration no higher than the Time Weighted Average (TWA) of the Threshold Level (TLV). The time-weighted exposure limit (TLV-TWA) is the average concentration for a normal working day exposure (40 hours per week, 8 hours per day) with no adverse effect. Table 3 lists the TLV-TWA for various refrigerants. In the following analyzes and examples, the TLV-TWA shall be used as the threshold for activating the ventilation. For chemicals for which a TLV-TWA is not available, 10% of the RCL is used as the detector threshold.

Table 3. Refrigerant quantities used to calculate aeration rate Refrigerant RCL (a) Concentration set point (b) ppm g/[m.sup.3] lb/Mcf ppm mg/[m.sup.3] (v/v) (v/v ) R -12 18,000 90 5.6 1,000 49,000 210 13 1,000 33 41,000 120 7.100 12,000r-36.600 700 700r-123 9.100 57 5.700r-14 10,000 56 1.5 1,000 5 5 , 5 5.5.600r -125 75,000 370 23 1,000 4.909r-134a 50,000 210 13 1,173r-1143a 21,000 70 4.5 1,000 32 2.0 1,200r-170 9 0.54 1,000 1.230r-245fa 34,000 190 12.400 19,000 R-2900R-245FA 5.300 10 0.56 1.804r-404a 130,000 31 1,000 3.92R-407C 76,000 270 270,000 39,000 39,000r-507a 130,000,043r-600A 4,000 10 0.6 400 960R-717 (c) 320 0 0.014 25 17R-1270 1,000 2 0.1 500 861refrigerant m [PHI] [Q.sub.max] lb/Mcf sec cfm L/sR-12 0.31 39 0, 33 44,400 21,000R-22 0.23 32 0.34 23.5 0.74 0.72 0.72 0 1R-1240 24 0.61 95,700 45,200R-32 0.49 28 0.34 73,600 34,800R-123 0.36 161 0.049 2,550 1.210R-124 0.35 51 0.3 48.600 23,000r-125 0.54 24.700R-14A 0.27 40 0.35 9.310r-143a 0.22 33 0.44.600r-1529 85.200 40.077 39 0.64 873.000 412,000r-245fa 1.2 980 1.690 R-290 0.4 372,000 176,000R-404A 0.5 14.600 6.860R-407C 1.7 0.4 22.200 27 0.42 18.640r-507a 0.26 6.820R-600a 0.06 87 0.29 164.000 77.200R-717 (c) 0.0011 45 0.21 10,100,000 4,730,000R-1270 0.054 48 0.4 2,310,000 1,090.0 ANSI/ASHRA-Aus00(a.ASHA) Aus00(a.ASHA) Standard 34-2010 (ASHRAE 2010). In the event of a dispute, Standard 34 shall prevail. (b.) TLV/TWA value or 1/10 RCL if TLV/TWA is not available (c.) R-717 see IIAR Standard 2 (IIAR 2008) for required ventilation rates

Refrigerant concentration in a room

A mass balance equation is used to solve the concentration in a space as a function of pollutant sources and losses, mainly ventilation. Several assumptions are made in deriving and solving this mass balance:

1. The room is well mixed. If the contaminating refrigerant is released into the enclosed space and is in gaseous form, the emitted mass is immediately diluted in the volume of the space. Without much more sophisticated tools to account for densities, mixing rates, and other parameters used to determine stratification and the concentration gradient, this assumption will underestimate concentration in one part of space and overestimate concentration in others. For example, if the refrigerant is heavier than air and the aeration rate is initially small (little mixing), the concentration would be higher at the bottom and lower near the ceiling. Because of this, many emergency exhaust intakes are located closer to the ground. Once the aeration rate is increased, better mixing will most likely occur.

2. The impurity is inert. There are no reactions in the volume of space, neither homogeneous nor heterogeneous.

3. There is no nucleation or condensation. Once the contaminating refrigerant is gaseous, no nucleation or condensation will occur.

For a room volume V, the change in concentration C is equal to the emission rate E minus the ventilation Q of the pollutant from the room:

VdC(t)/dt = E(t) - QC(t) (19)

Dividing all the terms by the volume gives a first-order differential equation.

dC(t)/dt + λC(t) = E(t)/V (20)

By multiplying all terms by [eλt] and applying the reverse chain rule, an intermediate solution for C(t) is found:

[e.sup.λt] dC(t)/dt + [e.sup.λt]C(t) = [e.sup.λt]E(t)/V (21 )

d/dt[C(t) [e.sup.λt]] = [e.sup.λt]E(t)/V (22)

C(t) = [e – λt][Integral[e λr]E(t)/V dt] (23)

Assuming a constant emission rate E(t) = [E0] and an initial condition of C(0) = [C0], the analytical solution for C(t) is:

C(t) = [E0]/Q(1-[e0]/Q(1-[e0-[lambde]t]) + [C0][e0-[lambde]t] (24 )

For most other emission rates, a numerical approach is required to find a solution.

The analytical solutions consist of two basic mechanisms. The drop in concentration due to aeration (the only loss accounted for in this approach) is given by the [C 0 ][e - λ t ] (24) term. The second term or terms containing the (1-[e-λ]) term or variants thereof represent the concentration's approximation to an anasymptote, which represents a steady-state concentration. This term is multiplied by the source emission to get the magnitude.

For each unique set of source and loss terms (emission rate and ventilation rate), the equation should be reviewed with a new C0 and E0 and any other parameters that may have changed. Each time emissions or ventilation rates change, the system tries to reach a new equilibrium. For example, for a short time at the beginning of the leak, the emergency suction has not started, so the concentration in the room increases. As the aeration rate increases, the steady-state concentration value decreases.

Assuming a constant rate of emission, a steady-state concentration will eventually be reached. Again, this represents the worst case scenario. At a constant emission rate, the concentration of refrigerant in the room is independent of the room volume and is only a function of the emission rate and the RCL, as shown below:

C(t) = [E0]/[lambde] = [E0]/Q (25)

Therefore, to calculate the venting rate for the worst case scenario (steady state for a given leak), the formula is:

Q = [E0]/RCL (26)

where RCL is the recommended concentration limit given by Standard34 and [E0] is determined by the accident scenario.

The steady-state solution is conservative as it only examines the maximum emission rate, while ignoring factors such as room volume, time-dependent emission rate, and variable aeration (i.e. low aeration when the concentration is below the TLV and high thereafter). can. By calculating the concentration as a function of time, these factors can be taken into account.

Although numerical integration of Equation 25 by various implicit or explicit methods would give a purely numerical solution, it can be used for any time step, given the value for the initial concentration, [C 0 ], and the emission rate, [E 0], correlates with each time step, n. The superscript n denotes the time step.

[Cn] = [En]/[Qn](1-[e.sup.[[lambde]n]([tn]-[ tn-1]])]) + [Cn-1] [e[[lambde]n]([tn] - [tn] n-1]])]) (27)

The release duration is defined by the total mass of the refrigerant. The emitted liquid, either vaporized or spilled as liquid and then vaporized, is limited by the reservoir of available refrigerant. Therefore, when liquid emission ceases, the flashed vapor ends, and the liquid pool stops growing and decreases in size as the coolant evaporates. Fig. 9 is an emission rate pattern diagram showing the total liquid emission rate, the flash emission rate, and the evaporative emission rate. The figure shows the relationship between the different mechanisms that contribute to the refrigerant vapor in the room.

[FIGURE 9 OMIT]

Evaporation from the pool creates difficult nonlinearities in the solution as it introduces additional variables that need to be resolved, such as B. the depth of the pool, the air speed at the surface of the pool and the size of the pool, which requires a floor area. Comparison a Leaks from a 0.25 inch (6.35 mm) hole with pool evaporation and from a 0.5 inch (12.7 mm) hole without evaporation gave similar gas emission rates for most refrigerants. Therefore, for the selected accident scenario, the simplification was made to increase the hole size but ignore evaporation from the liquid pool. It is also recommended that the room has a drain to drain liquids from the floor.

For each time step, the liquid emission rate is divided into the blinking part and the liquid pool that evaporates but is ignored. The concentration in the room is calculated from this total vapor emission rate. The maximum concentration is found and a routine solves for the aeration rate such that the maximum concentration does not exceed the RCL. Some of the variables that can be changed are: the total refrigerant mass, the hole diameter for the accident scenario, the temperature and pressure of the refrigerant in the tank, the volume of the space, and the RCL and TLV TWA values ​​(which depend on the refrigerant).

The steady-state solution gives the simplest calculation for the required aeration rate, but is also the most conservative. This can be a starting point for the calculation, but a time-varying analysis that takes into account that the emission rate does not reach equilibrium gives a much more accurate solution. The transient analysis ends by calculating the required ventilation rate using the concentration in the room at the maximum emission rate. Depending on a number of variables, including the mass of the refrigerant, the size of the room, and the manner in which the pool of liquid is contained, ventilation requirements may be higher than previously specified in Standard 15 (ASHRAE 2010) using the [cubic root of (100) ] shown. G equation (for lbs of refrigerant). With lower masses and larger spaces, the ventilation rate may be lower and the RCL may not even be reached.

Safe room volume

The safe volume of space, as defined herein, is the volume of space required to keep the leaked refrigerant concentration below the RCL if the entire mass has been leaked and immediately mixed. It also takes into account the flashed portion of the liquid release from the accident scenario. The safe space volume can be defined as follows:

[Vs] = G[PHI]/RCL (28)

The values ​​for [PHI] can be found in Table 3.

The safe space volume fraction f is defined as the ratio of the space volume to the safe space volume (V/[V s ]). This dimensionless variable includes refrigerant mass, room volume, RCL and accident scenario. If the safe space volume fraction f is greater than one, no additional ventilation is required if an accidental release occurs.

Verzögerungszeit des Detektor

The time between the onset of refrigerant release and the fan turning on can have a large impact on the ventilation rate required. Increased ventilation is required to slow down the increase in concentration in the room. When the supply of available refrigerant is exhausted, ventilation will continue to purge the room. An example of this effect on concentration is shown in Figure 10, which uses the same parameters for R-134a as the accident scenario. The concentration increases until all available refrigerant has escaped, after which ventilation purges the room of refrigerant. The goal is to keep the maximum concentration below the RCL at all times. If left too long, the concentration will exceed the RCL. Ventilation slows the rise in concentration and eventually flushes the room, but the RCL has already been exceeded.

The maximum lag time was found for which some degree of ventilation could prevent the concentration from exceeding the RCL. This aeration rate is too high. Figure 11 shows that delaying ventilation increases ventilation demand dramatically when the maximum delay time is reached. To limit the ventilation required, a detector delay time of 50% of the maximum detector delay time was provided. All maximum aeration rates ([Qmax]) are calculated at this point.

[FIGURE 10 OMIT]

[FIGURE 11 OMIT]

For each refrigerant and machine room, the allowable detector delay time in seconds is calculated using m (Table 3) and f:

[td] [less than or equal to] mf for f < 1 (29)

Recommended ventilation rate

By calculating the required ventilation rate as a function of the safe room volume fraction f, the required ventilation rate is fitted to the data (see Figure 12) on:

Q = 0 for f > 1(30)

Q = [Qmax](1 + 0.3f – 1.3[f2]) for f < 1(31)

The values ​​for [Qmax] are given in Table 3.

DISCUSSION

[FIGURE 12 OMIT]

It is useful to compare this method of calculating ventilation requirements with that of Standard 15 (ASHRAE 2010). Because the approach in this work includes many more variables than just refrigerant mass, the ventilation demand trend is not quite as predictable as the square root of mass. In general, the ventilation demand decreases when the mass of the refrigerant is lower, the room volume is larger, the refrigerant is less volatile and the RCL is higher. To compare the two methods, a comparison was calculated using cold rooms designed as shown in Table 4. In some cases, the need for ventilation increases. For others it decreases. For some spaces, the safe volume fraction is above unity, so no additional ventilation is required to prevent total mass leakage from exceeding the RCL in the space.

Table 4. Comparison of proposed ventilation requirements to current standard 15 Cold room Refrigerant Mass [PHI] RCL f [Qmax] Volume [ft.sup.3] lb lb/Mcf cfm3300 134a 124 0.35 13 0.99 19.8007956 134a 400 0.35 13 0.74 19.8004730 123 750 0.049 3.5 0.45 255033.895 123 1050 0.049 3.5 2.31 25505299 134a 355 0.35 13 0.55 19.80036.209 909036.209 19.800878888888814a 760 0,35 1300 0,59 19.800878888888814a 760 0,35 1300 0,59 19.800878888888814a 760 0,35 1300 0,59 19.8008788888888814a 760 0,35 1300 0,59.800878888788888814a 760 0,35 1300 0,59.800878888788888102. 13 1.02 19.80029.700 134a 625 0.35 13 1.77 19.800Chiller Room 17.603 275720.925 0 275729.700 0 2500

CONCLUSIONS

The current recommendation for ventilation in chiller rooms in Standard 15 only considers the mass of refrigerant in the room and has no strong scientific justification for its existence. This work presents a ventilation requirement for refrigeration machine rooms in the event of a release. The approach is simple enough to be codified or included in a standard and includes the main variables that would affect ventilation requirements: mass of refrigerant in the room, volume of the room , refrigerant properties, concentration limit and a reasonably conservative and realistic accident scenario. Further work is needed to extend the refrigerants with RCL values ​​and to better define some of the refrigerant properties in Standard 34 to complete this work for all refrigerants.

THANKS

This work was funded by ASHRAE Research Project 1448-RP through TC4.3 - Ventilation Requirements & Infiltration.

NOMENCLATUR

A = area

[c0] = discharge coefficient

[cp] = specific heat at constant pressure

[cv] ​​= specific heat at constant volume

d = diameter

E = mass flow

[ff] = friction factor

f = safe volume fraction, V/[V S ]

G = Masse

g = acceleration due to gravity

H = height

[H.sub.vap] = enthalpy of vaporization (difference between [H.sub.] and [H.sub.t] at boiling point and ambient pressure)

p = pressure of the liquid in the container,

K = loss of fit coefficient

k = specific heat ratio [c.cub.p]/[c.sub.v]

L = Length

M = Masse

[Mw] = molecular weight

m = detector delay time influence factor (Table 3)

Re = Reynolds number

Q = flow rate

R = universelle Gaskonstante

S = Entropy

[DELTA][Svap] = entropy of vaporization (difference between [Sv] and [St] at boiling point and ambient pressure)

T = Temperature

t = time

U = air velocity

V = Volume

[Vp] = vapor pressure (at [T2])

[Vph] = vapor pressure of hydrazine (at [T2])

[Vs] = safe volume of space, M[PHI]/RCL

Greek

[epsilon] = roughness height

[lambde] = air change rate, Q/V

v = kinematic viscosity

[THETA] = Angle relative to the upward vertical direction formed by the liquid surface remaining around the perimeter of the tank

[rho] = density

[PHI] = percentage of flashing liquid found in Table 3

indices

b = boil

d = detector delay

f = blinks

h = Loch

l = liquid

v = steam

s = surface

0 = ambient or initial value

1,2,3 = states

max = maximal

DISCUSSION

Joy Kohler, Engineering Manager, Johnson Controls, York, PA: The report cites just one example of fluorocarbon leaks in engine rooms: a fatal accident in Alaska. Most of the data comes from ammonia systems, which differ significantly from fluorocarbon systems due to the self-alarming nature of ammonia and the typical construction of ammonia systems. What efforts have been made to find such data? Accident data must certainly be available for such systems.

Scot K. Waye: The main databases searched for engine room refrigerant leaks are the following: MARS, NRC, ARIP, OSHA and NIOSH. Table 1 of the paper lists the number of accidents reported to the NRC for different refrigerants. The number of spills of fluorocarbon refrigerants reported for the given period is very small, especially when compared to ammonia. The same applies to the OSHA database. No accidents involving fluorocarbon refrigerants were found in the MARS database and only one incident (the Alaskan incident) was found in the NIOSH database. Summaries of incidents reported in the NRC database are attached to the full report for ASHRAE RP-1448. [Editor's note: Final reports of ASHRAE research projects are available for members to download free of charge at www.ashrae.org.] Efforts were made to locate additional databases and sources for additional reported leaks involving non-ammonia refrigerants, however these provided no additional incidents.

For many refrigerants, apart from ammonia, the release level that triggers reporting requirements is high, which we believe leads to underreporting of fluorocarbon leak incidents. Some refrigerants do not have a reportable quantity listed at all. The report of the Alaska incident was due to one fatality, leading to an investigation and report. For some cases it is believed that in the absence of injury or death, reporting the leak was not mandatory under current requirements; this also leads to under-reporting. We recognize that better reporting and more data depicting fluorocarbon refrigerant leaks would improve understanding of the issue, but this data was not available to us through public and open sources. Although much of the data was for ammonia from actual releases, trends from potential leak locations and sources can be examined. In addition, a threshold probability for exceeding the emission rate can be estimated for worst-case emission scenarios.

REFERENCES

AIChE (American Institute of Chemical Engineers). 1996. Guidelines of Vapor Cloud Dispersion Models, 2nd edition. New York: Center for Chemical Process Safety.

API (American Petroleum Institute). 1996. Manual for ModelingHypothetical Accidental Releases to the Atmosphere. API-VERÖFFENTLICHUNG 4628.

ASHRAE. 2009. 2009 ASHRAE-Handbuch – Grundlegend. Atlanta: AmericanSociety of Heating, Air-Conditioning and Refrigeration Engineers, Inc.

ASHRAE. 2010. ANSI/ASHRAE Standard 15-2010, Safety Standard for Refrigeration Systems. Atlanta: American Society of Heating, Air-Conditioning and Refrigeration Engineers, Inc.

ASHRAE. 2010. ANSI/ASHRAE Standard 34-2010, Refrigerant Designation and Safety Classification. Atlanta: American Society of Heating, Air-Conditioning and Refrigeration Engineers, Inc.

Brown, R. 2005. Engine Room Ventilation for Industrial Refrigeration Systems: A Rational Engineering Approach. Technical paper no. 5th IIAR Ammonia Refrigeration Conference and Exhibition. Acapulco, Mexico.

Fenton, D.L. and W.V. Richards. 2003. User Guide for ANSI/ASHRAE Standard 15-2001 Safety Standard for Refrigeration Systems. ASHRAE Special Project SP-93. Atlanta: American Society of Heating, Air-Conditioning and Refrigeration Engineers, Inc.

IIAR. 2008. ANSI/IIAR 2-2008 (Anhang A), Americanation NationalStandard for Equipment, Design and Installation of Closed-CircuitAmmonia Mechanical Refrigerating Systems. Alexandria: International Institute of Ammoniak Refrigeration.

MAHB (Bureau for Major Accident Hazards). 2001. eMARS – Major Accident Reporting System. www.emars.jrc.ec.europa.eu/?id=4.

Moran, M.J. and H.N. Shapiro. 2000. Fundamentals of Engineering Thermodynamics, 4th ed. New York: John Wiley and Sons.

NIOSH (National Institute for Occupational Safety and Health).2010. Fatality Assessment and Control Evaluation (FACE) Program. www.cdc.gov/niosh/face/.

NRC (National Response Center). 2010.www.nrc.uscg.mil/download.html

OSHA (Occupational Safety and Health Administration). 1992. Process Safety Management (PSM) of Highly Hazardous Chemicals Standard.29 CFR 1910.119.

OSHA (Occupational Safety and Health Administration). 2010. Fatality and Catastrophe Investigation Summaries. www.osha.gov/pls/imis/accidentsearch.html

Richards, W.V. 1986. How codes and regulations affect your facility design. IIAR Ammonia Refrigeration Conference and Expo.Innisbrook, FL.

Sallet, D.W. 1990. Critical two-phase mass flow rates of liquefiedgases. Journal of Loss Prevention in the Process Industries 3:38-42.

Seidl, R. and S.T. Taylor. 2005. Exhaust system sizing for chiller rooms. Taylor Engineering, Alameda, CA.

Stoecker, W. F. 1998. Industrial Refrigeration Handbook. New York: McGraw-Hill Companies, Inc.

US EPA. 1992. Workbook of Screening Techniques for AssessingImpacts of Toxic Air Pollutants (Revised). Office of Air QualityPlanning and Standards.

U.S. EPA. 2001. Consolidated List of Chemicals Subject to the Emergency Planning and Community Right-to-Know Act (EPCRA) and Section 112[R] of the Clean Air Act. EPA 500-B-01-003. Office of Solid Waste and Emergency Response (5104).

US EPA, USA. US EPA. 2010. www.epa.gov/emergencies/tools.htm#arip

Wilson, D.M. 1981. Longitudinal wind diffusion and source transients. Atmospheric Environment, 15:489-495.

This paper is based on findings from ASHRAE research project RP-1448.

Scot K. Waye, PhD, PE member ASHRAE

Ronald L. Petersen, PhD member ASHRAE

Anke Beyer-Lout

Scot K. Waye is chief engineer, Ronald L. Petersen is director and vice president, and Anke Beyer-Lout is project scientist at CPP, Inc. of Fort Collins, CO.