A national laboratory of the U.S. Department of Energy

Office of Energy Efficiency & Renewable Energy

National Renewable Energy Laboratory 

Innovation for Our Energy Future 

Analysis of Buoyancy-Driven 

Ventilation of Hydrogen from 

Buildings 

Preprint 

C.D. Barley, K. Gawlik, J. Ohi, and R. Hewett 

National Renewable Energy Laboratory 

To be presented at the 2nd International Conference on 

Hydrogen Safety 

San Sebastian, Spain 

September 11–13, 2007 

Conference Paper 

NREL/CP-550-41081 

August 2007 

NREL is operated by Midwest Research Institute ● Battelle Contract No. DE-AC36-99-GO10337 
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ANALYSIS OF BUOYANCY-DRIVEN VENTILATION OF HYDROGEN 

FROM BUILDINGS*

Barley, C.D.1

, Gawlik, K.2, Ohi, J., and Hewett, R. 

National Renewable Energy Laboratory, 

1617 Cole Boulevard, Golden, Colorado 80401, USA 

1

dennis_barley@nrel.gov 2keith_gawlik@nrel.gov

ABSTRACT

When hydrogen gas is used or stored within a building, as with a hydrogen-powered vehicle parked in 

a residential garage, any leakage of unignited H2 will mix with indoor air and may form a flammable 

mixture. One approach to safety engineering relies on buoyancy-driven, passive ventilation of H2

from the building through vents to the outside. To discover relationships between design variables, 

we combine two types of analysis: (1) a simplified, 1-D, steady-state analysis of buoyancy-driven 

ventilation and (2) CFD modeling, using FLUENT 6.3. The simplified model yields a closed-form 

expression relating the H2 concentration to vent area, height, and discharge coefficient; leakage rate; 

and a stratification factor. The CFD modeling includes 3-D geometry; H2 cloud formation; diffusion, 

momentum, convection, and thermal effects; and transient response. We modeled a typical residential 

two-car garage, with 5 kg of H2 stored in a fuel tank; leakage rates of 5.9 to 82 L/min (tank discharge 

times of 12 hours to 1 week); a variety of vent sizes and heights; and both isothermal and nonisothermal conditions. This modeling indicates a range of the stratification factor needed to apply the 

simplified model for vent sizing, as well as a more complete understanding of the dynamics of H2

movement within the building. A significant thermal effect occurs when outdoor temperature is 

higher than indoor temperature, so that thermocirculation opposes the buoyancy-driven ventilation of 

H2. This circumstance leads to higher concentrations of H2 in the building, relative to an isothermal 

case. In an unconditioned space, such as a residential garage, this effect depends on the thermal 

coupling of indoor air to outdoor air, the ground (under a concrete slab floor), and an adjacent 

conditioned space, in addition to temperatures. We use CFD modeling to explore the magnitude of 

this effect under rather extreme conditions. 

NOMENCLATURE 

c = Concentration of H2, by volume (dimensionless, 0-1 in formulas, 0%-100% in graphs) 

g = Acceleration of gravity = 9.81 m/s2

h = Height between vents, m 

A = Vent area (top = bottom), m2

D = Vent discharge coefficient (dimensionless, 0-1) 

D*

 = Apparent value of D based on CFD results (dimensionless, 0-1) 

F = Vent sizing factor (dimensionless) 

NTP = Normal temperature and pressure (20°C and 1 atm) 

P = Total pressure (Pa) 

Q = Volumetric flow rate through a vent (m3

/s) 

S = Source rate of H2 (leak rate), m3

/s 

Δ = Difference 

δ = Ratio of densities of H2/air at NTP = 0.0717 (dimensionless) 

ρ = Density (kg/m3

) 

φ = Stratification factor = cT/cavg (dimensionless) 

__________________________ 

*

 This manuscript has been authored by Midwest Research Institute under Contract No. DE-AC36-99GO10337 with the U.S. Department of 

Energy. The United States Government retains and the publisher, by accepting the article for publication, acknowledges that the United 

States Government retains a non-exclusive, paid-up, irrevocable, world-wide license to publish or reproduce the published form of this 

manuscript, or allow others to do so, for United States Government purposes.

1
Subscripts: 

1,2,3,4 Locations shown in Fig. 1 (Section 2) 

air Properties of air 

avg Spatial average over the path 3-4 in Fig. 1 (Section 2) 

o Outdoor 

B Bottom vent 

T Top vent 

1.0 INTRODUCTION

When hydrogen gas (H2) is used or stored within a building, as with a hydrogen-powered vehicle 

parked in a residential garage or commercial service facility, any leakage of unignited H2 will mix 

with indoor air and can potentially form a flammable mixture. The general approach to fire safety is 

to avoid a combination of fuel, an oxidizer, and heat (or spark) that completes the “fire triangle” and 

enables combustion. Oxygen is unavoidably present in occupied buildings. Also, the static electricity 

that is routinely generated by the movement of occupants is sufficient to ignite a mixture of H2 and air 

[Ref. 1, p. 11]. Therefore, the only practical approach to safety in this context is to control the fuel so 

that a flammable mixture does not occur. The most widely accepted flammability range is 4.1% to 

75% H2 by volume [1]. To provide a safety margin, many standards [e.g. 2,3] specify a safe limit of 

25% of the lower flammability limit (LFL), which for H2 amounts to 1% by volume. Swain et al. 

[4,5] have measured the LFL in H2 leakage jets between 6.6% and 8.1%. However, the current study 

involves H2-air mixtures that are practically stationary, and we have no additional information 

regarding the relationship between velocity and LFL. 

The well-known approaches to indoor air quality problems, of which this is an example, include (1) 

source control, (2) air treatment, (3) natural (passive) ventilation, and (4) mechanical ventilation. 

Source control involves preventing the infusion of a pollutant, H2 in this case, into the indoor air. One 

method is to keep the fuel system outside the building. An automobile can be parked in a driveway or 

carport. Although this may be acceptable to some people, the use of residential garages is very well 

established and would be difficult to eliminate. Another source control method is quality control in 

the manufacture of vehicles, to limit the rate of fuel leakage, as in Standard SAE J2578 [6]. However, 

higher leakage rates must also be considered, as these may result from faulty repair work or collision 

damage after the car leaves the factory. Air treatment, or cleaning, involves the removal of a 

pollutant from indoor air. It may be possible to capture leaked H2 within a building in a chemical 

process with either a mechanical or a passive apparatus. For example, H2 does react with halogens. 

This is an area for possible future research. Concerns include reliability, response time, saturation 

capacity, reactant standby life, and cost. 

In this paper we describe our analysis of buoyancy-driven, passive ventilation of H2 from buildings 

through vents to the outside. Our goal is to ascertain the relationship between vent design, leakage 

rate, maximum concentration, and other variables, which leads to design guidelines, an understanding 

of the limitations of this approach, and recommendations for codes and standards. We pursue this 

goal by combining two types of analysis: (1) a simplified, one-dimensional, steady-state analysis of 

buoyancy-driven ventilation; and (2) computational fluid dynamics (CFD) modeling. The simplified 

model enables an algebraic solution that yields a closed-form expression relating the concentration of 

H2 to vent size, height, and discharge coefficient; leakage rate; and a stratification factor. However, 

the stratification factor and vent discharge coefficient depend on dynamics that are beyond the scope 

of this model. The CFD modeling provides additional detail and indicates a range of the stratification 

factor needed to apply the simplified model. Because of the limitations of passive ventilation shown 

by this study, we also address mechanical ventilation rates. 

The range and frequencies of occurrence of leakage rates that will occur with H2 vehicles are 

unknown to us, despite our literature search. In an effort to bracket a range of possible leakage rates 

2
and study sensitivities to this variable, we modeled rates ranging from 5.9 to 82 L/min, which 

correspond to leak-down times of 12 hours to 7 days for a 5-kg tank of H2. We also calculate required 

mechanical ventilation rates over a wider range: 1.4 to 166 L/min, or 6 hours to 29 days. The 29-day 

leak-down time is based on a safety standard for vehicle manufacture [6]. We selected the 6-hour 

leakage rate arbitrarily, to provide an ample range for sensitivity study. 

The TISEC Sourcebook for Hydrogen Applications [1] presents a compilation of basic information 

about H2 properties, safety engineering, codes, and standards. Cadwallader and Herring [7] present 

similar general information as well as historical accounts of accidents involving H2. Swain and Swain 

[8] have modeled transient accumulation rates of H2 within a cloud above an H2 leak in a passively 

ventilated room, with leak rates of 10 to 1000 L/min. Breitung et al. [9] have modeled transient H2

cloud formation resulting from pulsed release of small amounts of H2 (i.e., insufficient amounts to 

form a combustible concentration if fully mixed in the enclosure). Swain et al. have used helium to 

study H2 leakage scenarios [10] and have tested H2 ignition and combustion scenarios [4,5,11]. 

Papanikolaou and Venetsanos [12] have modeled three of the helium experiments performed by 

Swain et al. The present study addresses steady-state concentrations of H2 resulting from sustained 

slow leaks; modeling of thermal effects due to high outdoor temperature; and passive vent sizing as a 

function of leakage rate and other variables. 

2.0 SIMPLIFIED MODEL 

The concept for the simplified model arose from the initial results of our CFD modeling, which is 

described in Section 3. Fig. 5 (Section 3) shows a typical steady-state H2 stratification pattern 

computed by the CFD model. The concentration varies significantly in the vertical direction, but 

much less in the horizontal directions (except for the plume rising from the leak site). This suggests 

that a one-dimensional, steady-state analysis might capture the basic dynamics of this process and 

provide insights into relationships between variables. Indeed, the simplified model has proven useful 

in this manner. Previous researchers have conducted one-dimensional, steady-state analyses of 

thermally driven airflow through vents in vertical walls [13,14,15]. However, our literature search 

did not reveal any prior formulation relating gas concentration to leakage rate and vent size for 

ventilation driven by a buoyant gas. 

Figure 1. Illustration of the simplified model. 

3
As shown in Fig. 1, we consider a source of H2 gas (the leak) within an enclosure (such as a garage) 

with two vents to the outside, near the top and bottom of a side wall. To formulate the relationship 

between the buoyancy pressure and the pressure drops across the vents, we draw a closed contour (1-

2-3-4) through both vents. Summing the total pressure differences around the loop, we write: 

ΔP1-2 + ΔP2-3 + ΔP3-4 + ΔP4-1 = 0 (1) 

The two vertical segments of the loop (1-2 and 3-4) represent the buoyancy pressure as the difference 

between the weights of the inside and outside air columns. Assuming isothermal conditions between 

the garage interior and the outside air (thermal effects are discussed in Section 4): 

ΔP1-2 + ΔP3-4 = g h ρair cavg (1-δ) (2) 

The relationship between pressure drop and airflow through the top vent is [16]: 

T

T

P ADQ

ρ

2Δ −41 = (3) 

A similar equation applies to the bottom vent. The continuity equation for air and H2 is: 

QT = QB + S (4) 

The continuity equation for H2 alone is: 

QT cT = S (5) 

Combining Eqns. 1 to 5 and reducing to a non-dimensional form results in this isothermal vent-sizing 

equation for buoyancy-driven ventilation: 

( )( )

( )

2

1

3

2

2

1

1

111 2

⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

−

−+−− ≡ =

T

T T

c

c c gh S

AD F

δ

δ φ (6) 

The term involving the vent area, A, defined as the vent sizing factor, F, may also be thought of as the 

dimensionless vent area. It implies that the required vent area is proportional to the leakage rate and 

inversely proportional to the square root of the height between the vents, if other factors are constant. 

The right-hand term, involving the H2 concentration, is more complex, because this factor enters the 

analysis in three places. The stratification factor, φ, occurs because the buoyancy force depends on 

the average H2 concentration over the height (cavg in Eqn. 2) whereas the continuity equation (Eqn. 5) 

involves the concentration at the top vent, cT. (The density at the top vent also depends on cT, in Eqn. 

3.) It may seem counterintuitive that larger vents are required when the H2 is more stratified. This 

relationship occurs because, given the same concentration at the top vent, there is less buoyancy in the 

indoor air column when the H2 is less distributed over the height. The isothermal vent sizing equation 

(Eqn. 6), which is illustrated by the curves in Fig. 2, is useful for vent design purposes. However, 

values of the stratification factor, φ, and the vent discharge coefficient, D, are needed, and these are 

pursued in the following section. 

3.0 CFD MODELING, ISOTHERMAL 

CFD modeling provides detail and accuracy (pending experimental verification) beyond the 

capabilities of the simplified model and leads to a more thorough understanding of H2 movement in 

the building. It includes 3-D geometry; diffusion, momentum, convection, and thermal effects; H2

4
cloud tracking; and transient response. It can also indicate a range of the stratification factor needed 

to apply the simplified model for vent sizing. 

0

1000

2000

3000

4000

5000

6000

7000

0.25 0.50 0.75 1.00 1.25 1.50 1.75 2.00 2.25 2.50 2.75 3.00

cT = H2 Concentration by Volume, %

F = Vent Sizing Factor

1

2

4 5 6 7

3

φ = 1.0, 1.5, 2.0, 3.0 left to right

Figure 2. Comparison of simplified model (curves are based on Eqn. 6) 

and isothermal CFD results (points 1 to 7, Section 3) 

. 

Figure 3. Home used for isothermal CFD case study (Cases 1-7), built by Pulte Homes 

in Las Vegas, Nevada. Courtesy of the Building America Program [17]. 

For our initial case study we selected a typical home, shown in Fig. 3, which features a two-car garage 

with an A-frame roof. The garage floor is 6.40 m wide by 6.71 m deep. The roof slope is 5/12, and 

the roof ridge is 4.06 m above the floor. The garage volume is 146 m3

. In this case study we assume 

the garage is an isolated zone with no transfer of H2 into the remainder of the house. The assumed 

leak source is a vehicle fuel tank containing 5 kg of pressurized H2 gas. This amount of H2 occupies a 

volume of 59.7 m3

 at NTP, which amounts to 41% of the garage volume; so a uniform mixture of 10 

5
times the LFL is possible. Tank pressures of 340 to 680 atm (34.5 to 68.9 MPa) are to be expected, 

although the pressure does not enter our computations directly. We modeled leakage rates ranging 

from 5.9 to 83 L/min, corresponding to tank leak-down times of 12 hours to 1 week. To reduce the 

required computational effort, we set up our model with bilateral symmetry, so we model half of the 

garage and the results represent each half. The leak is located 1.07 m above the floor, approximating 

the height of a leak from a vehicle, under the roof ridge, halfway between the front and back of the 

garage. The rectangular vents are located in the front wall, under the roof ridge, one adjacent to the 

floor (in the garage door) and the other as near to the roof ridge as we can locate it in the triangular 

gable. 

We used the FLUENT version 6.3.26 CFD software [18]. We used the poly mesh option for 

computational economy, with first order discretization of convection and time for speed and stability. 

A sensitivity study showed that the solution is insensitive to grid density at about 40,000 cells. 

Evidently, this grid density is sufficient to minimize the effects of numerical diffusion.1

 We then 

used this grid density for our further computations. The average size of a grid cell is about 1.8 L, 

although we used a higher mesh density near the leak and the vents where gradients are high. We 

used both the renormalization group (RNG) k-ε turbulence model with the differential viscosity 

option enabled and the laminar model. The results indicate turbulence intensities lower than 1% in 

more than 85% of the domain in all the cases we ran, and more than 90% of the domain in four of the 

cases. Other researchers have found that k-ε turbulence models tend to overpredict mixing of gas 

released from slow leaks [12], thus underpredicting stratification, and the use of a laminar model will 

lead to a more stratified hydrogen distribution [9]. Based on Eqn. 6 and Fig. 2, higher stratification 

corresponds to higher H2 concentrations for a given vent size. Thus, the laminar model is more 

conservative for our purpose of safety engineering, so we proceeded to use the laminar model for the 

results we are reporting here. We modeled the inlet and outlet vents as pressure boundary conditions. 

The model determines static pressure differences due to hydrostatic head outside of the inlet and 

outlet vents, based on the specified ambient outdoor air density corresponding to the simulation 

conditions. The incompressible ideal gas law is used to determine indoor density as a function of 

temperature and the air-hydrogen mixture. Because the boundary of our grid corresponds to the 

garage envelope, we found it necessary to disable diffusion across the boundary in order to eliminate 

spurious H2 diffusion across the abrupt concentration gradient at the outlet vent, which would not 

occur in reality. Time steps were changed during the course of the simulation to speed convergence. 

Initially, the time step was 0.1 seconds to capture the transient behavior as the leak started. When the 

iterations to convergence at each time step dropped to below 10, the time step was doubled. The 

largest time step used was 0.4 seconds. Data and image files were saved every 10 minutes of flow 

me. 

odeling described in this section does not include thermal 

ffects, which are described in Section 4. 

 

ti

An actual leak from a pressurized tank might occur at a high velocity in any direction. On the other 

hand, if a leak occurs within the body of the car, H2 may emanate from the car at a low speed. We 

modeled the source of the leak as an 8-cm sphere from which H2 emanates at the lowest possible 

speed, driven only by displacement. We believe this is a worst-case scenario, because our modeling 

has shown that (1) momentum effects tend to augment buoyancy in driving vent flow and lead to 

lower H2 concentrations, and (2) momentum tends to destratify the H2-air mixture, which also leads to 

lower concentrations (see Eqn. 6 and Fig. 2). More modeling is needed to determine more 

specifically the effects of leakage velocity. In addition to plotting the concentration throughout the 

garage periodically, we also monitored the concentration at the outlet vent to determine when steadystate conditions were established. The m

e

1 Numerical diffusion is a consequence of discretizing into finite-difference equations the continuous NavierStokes equations of fluid motion. The discretized equations tend to exhibit greater diffusion than the continuous 

equations. One method to minimize numerical diffusion is to increase grid density until the converged solution 

does not change. Another method is to use higher order discretization schemes, but they can introduce 

instability in the solution process. 

6
In order to compare results of our CFD model with the simplified model, we ran the series of cases 

listed in Table 1 and plotted as points 1 to 7 in Fig. 2. In each case, the computation was continued 

until a steady-state condition was observed. Fig. 4 (“No Thermal Effects” curve) shows a typical 

transient response of the concentration at the top vent, although the time to steady state and the final 

concentration vary from case to case. Fig. 5 shows a typical stratification pattern under steady-state 

conditions, and Fig. 6 shows corresponding vertical concentration profiles at four locations in the 

arage. 

Table 1. Series of CFD cases plotted in Fig. 2. 

Figure 4. Develop vent for Case 8, 

showing the trend to steady state. 

1234567

Leak-Down Time, hr/5 kg 168 72 48 24 24 24 12

Vent Size, cm2 788 788 788 788 788 788 1576

Vent Offset, cm 0.0 0.0 0.0 0.0 15.2 30.5 0.0

Vent Height, m 3.650 3.650 3.650 3.650 3.345 3.040 3.599

H2 Conc. at top vent, % Vol. 0.47 0.79 1.04 1.55 1.63 1.69 1.75

Straification Factor (φ) 1.65 1.67 1.67 1.52 1.58 1.59 1.88

Discharge Coeff. (D*) 0.952 0.952 0.952 0.965 0.948 0.944 0.903

Specifications, CFD Case

Results

g

0.0

0.5

1.0

1.5

2.0

2.5

0.0 1.0 2.0 3.0 4.0 5.0 6.0 7.0 8.0

Time (hr)

H2 Concentration, % by Volume

No Thermal Effects

With Thermal Effects

ment of H2 concentration at the top

7
F 

Color scale is H2 mole fract alent to 2% H2 by volume. 

Elapsed time is 3 hours (steady state). 

Figure 6. Vertic tions in garage, 

corresponding to Fig. 5 and Case 7. 

e H2-air mixture and flow rates through the vents 

dicated by the CFD model. From Eqns. 1 and 3, 

igure 5. Graphic result of CFD modeling for Case 7. Left half of garage in Fig. 3 is shown here. 

ion; the full-scale value of 0.02 is equiv

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

4.0

4.5

0.000 0.004 0.008 0.012 0.016 0.020

Mole fraction

Height (m)

Near Vents

Back Corner

Back Wall Center

Side Wall Center

al H2 concentration profiles at 4 loca

We modeled the vents as open rectangles with no thickness and an aspect ratio of 4:3. As part of the 

comparison between the simplified model and the CFD results, we calculated values of the apparent 

vent discharge coefficient based on the density of th

in

8
2

1

2

2 2

)(2 ⎥

⎥

⎦

⎤

⎢

⎢

⎣

⎡

−

+ =

avgo

BoTT

ghA

QQ D

ρρ

ρρ (7) 

This formulation assumes equal values of D for the top and bottom vents, and it assumes that the flow 

through the vents is driven only by the static pressure difference. However, airflow in the CFD model 

may also be driven by momentum generated in the buoyant rise of H2 from the leak site. We use the 

symbol D*

 to denote the apparent value of D given by Eqn. 7 with QT, QB, ρT, and ρavg as computed by 

the CFD model. Values of D*

 determined by this method range from 0.903 to 0.965 (Table 1). This 

range is significantly higher than published values of D for similar configurations. Resistance 

coefficients reported by Idelchik [19] for orifices in a thin wall with infinite surface area correspond 

to a value of about D = 0.60. Building Science Consulting [20] cites a value of D = 0.61 for sizing 

transfer grilles for return air flows in residential forced-air systems. American Lung Association 

guidelines for sizing transfer grilles correspond to D = 0.72 [21]. Possible reasons for the difference 

between the published values of D and our values of D*

 include (1) the significance of momentum 

effects in driving airflow through the vents; (2) differences in the flow regime, indicated by the 

Reynolds number,2

 between our scenario and the literature references; (3) differences in geometry; 

and (4) CFD modeling issues. Further study, including model validation, is needed to understand this 

result. In the meantime, based on the literature references, the value D = 0.60 seems to be a 

reasonably conservative value for design purposes. In Cases 5 and 6, compared to Case 4, we 

explored the effect of offsetting the bottom vent above the floor and the top vent further from the roof 

ridge. The value of D*

 decreases slightly with increasing offset (Table 1), although the effect does not 

appear to be very significant. 

Values of the stratification factor, φ, range from 1.52 to 1.88 (Table 1). The agreement between the 

simplified model and the CFD results is indicated in Fig. 2 by points 1 to 7 lying between the solid 

curves for φ = 1.5 and 2.0. However, some differences between the two models (such as momentum 

effects) are absorbed in the values of D*

 that were calculated as described above and used in plotting 

points 1 to 7. Based on these results, the value φ = 2.0 appears to be reasonably conservative, 

pending model validation and further case studies. 

4.0 CFD MODELING, THERMAL EFFECTS 

Both the simplified model described in Section 2 and the CFD modeling described in Section 3 are 

based on assumed isothermal conditions. In this section we consider two thermal effects that may 

influence passive ventilation of H2. 

4.1 Expansion Heating 

Ordinarily, when many real gases expand from a higher to a lower pressure at normal temperatures 

(e.g. 293 K), their temperatures decrease, such as when air is vented from an automobile tire. 

However, because the Joule-Thomson inversion temperature for H2 is 202 K [22], the temperature of 

H2 may increase upon expansion, depending on the leak velocity. A maximum possible temperature 

rise of 30.1 K would occur if the leak occurred in both an adiabatic and an isenthalpic manner, as we 

have modeled it. Although this leakage scenario is not realistic, it does indicate the limiting case of 

maximal expansion heating. Based on the ideal gas law, this absolute temperature rise of about 10% 

corresponds to a density decrease of about 10%. However, based on the difference between the 

densities of air and H2, the buoyancy force is affected by less than 1%. This rules out any significant 

benefit of additional buoyancy due to expansion heating. In our modeling of the garage shown in Fig. 

 

2

 For cases 1-7, the Reynolds number (Re) for airflow through the top vent ranges from 4,700 to 9,100. The 

Idelchik reference is based on Re ≥ 10,000. The other literature references do not cite values of Re. 

9
3, the effect of including expansion heating in our model is negligible compared to the other effects 

we are studying. 

4.2 Reverse Thermocirculation 

A more significant thermal effect occurs when the outdoor air temperature is higher than the garage 

air temperature. Thermocirculation (the stack effect) opposes the buoyancy-driven ventilation of H2, 

and warmer outside air may enter the top vent while H2 accumulates near the ceiling. At the same 

time, the garage temperature is affected by the inflow of outside air as well as the thermal coupling 

between the garage and (1) the conditioned space in the home, and (2) the ground temperature, as 

through a concrete slab floor, and (3) the outdoor air by conduction through the envelope surfaces. 

Transient effects related to the thermal mass of the garage may also affect the garage temperature, 

although we did not model these effects. Such an outdoor-indoor temperature difference may occur 

on a hot day, with a garage that has strong thermal coupling to a low ground temperature and an airconditioned space and a low thermal coupling to the ambient air by conduction. Alternatively, a large 

temperature difference may occur on a cold morning, if the ambient air warms faster than a cold 

garage with significant thermal mass. The resulting H2 concentration can be significantly higher than 

that predicted by the isothermal models. 

In order to study the potential magnitude of this effect, we modeled a rather extreme hot-day scenario. 

We selected a home, shown in Fig. 7, with strong thermal coupling from the garage to the conditioned 

space (80.3 W/°C) and to the ground (32.7 W/°C) and weak thermal coupling to the outside air by 

conduction (16.4 W/°C). We assumed a high outdoor temperature (40.6°C), a low ground 

temperature (10°C), and a low air-conditioning setpoint (21.1°C). This garage is 6.40 m wide, 7.24 m 

deep, and 2.59 m high. We modeled the ambient temperature as a constant to explore steady-state 

effects. However, in reality, the ambient temperature varies with time, and thermal mass effects, 

which we did not model, are important. We modeled three cases with significantly different results. 

In Case 8 we used the 24-hour leak rate with 0.121 m2

 vents based on the isothermal vent sizing 

equation with cT = 2%, φ = 2.0 (so F = 723), and D = 0.60. As shown in Fig. 8, outside air initially 

flows in the top vent, causing H2 to accumulate in the top rear area of the garage. After about 1.7 

hours, the H2 buoyancy overcomes the reverse thermocirculation and begins escaping through the top 

vent. Fig. 4 shows the transient H2 concentration at the top vent compared to the same case with 

thermal effects disabled. The thermal case exceeds the design limit of 2% H2, (maximum 2.3%) 

whereas the isothermal case does not. The steady-state, volume-average garage air temperature is 

37.8°C in this case. 

Figure 7. Home used for CFD thermal case study (Cases 8 to 10), built by Heartland Homes in 

Pittsburgh, Pennsylvania. Courtesy of the Building America Program [17]. 

10
In Case 9 we used the 7-day leak rate with 494 cm2

 vents based on the isothermal vent sizing equation 

with cT = 1%, φ = 2.0 (so F = 2061), and D = 0.60. Because of the smaller vents, the garage remains 

cooler in this case than in Case 8 (36.4°C in the steady state). Also, because of the lower leak rate, 

the H2 destratifies more before reaching a high concentration. In the steady-state condition shown in 

Fig. 9, H2 is escaping through the bottom vent while reverse thermocirculation continues. The 

maximum H2 concentration in the garage, away from the plume, is about 1%, which is the same as the 

isothermal design value, although this agreement is considered coincidental. 

Figure 8. Graphic result of CFD modeling for thermal Case 8. Left half of garage in Fig. 7 is shown 

here. Color scale is H2 mole fraction; the full-scale value of 0.02 is equivalent to 2% H2 by volume. 

Elapsed time is 1.6 hours (transient response). 

Figure 9. Graphic result of CFD modeling for thermal Case 9. Left half of garage in Fig. 7 is shown 

here. Color scale is H2 mole fraction; the full-scale value of 0.015 is equivalent to 1.5% H2 by 

volume. Elapsed time is 31 hours (steady state). 

In an effort to identify a “perfect storm” scenario, we modeled an additional leakage rate between the 

previous two cases with opposite vent flow directions in the steady state. In Case 10 we used the 3-

day leak rate with 405 cm2

 vents based on the isothermal vent sizing equation with cT = 2%, φ = 2.0 

11
(so F = 723), and D = 0.60. In the steady state, H2 escapes through the top vent, with a maximum 

concentration of 5% away from the plume. This is the most significant thermal effect in the cases we 

have studied to date, with the H2 concentration reaching 2.5 times the value predicted by the 

isothermal vent sizing equation. However, this is not necessarily the worst-case scenario. 

Based on Cases 8 to 10, high outdoor temperature may or may not increase the H2 concentration 

relative to an isothermal case. As a safety engineering concern, this is a risk factor. Further study is 

needed to ascertain the maximum increase in H2 concentration that may result from reverse 

thermocirculation. 

5.0 WIND EFFECTS 

Wind pressure on a building may affect the buoyancy-driven ventilation of H2. Depending on the 

direction of the wind relative to the locations of the inlet and outlet vents, the wind may either help or 

hinder the ventilation. Under some conditions, a vent design that would otherwise limit the H2

concentration to the safe limit might be rendered ineffective because of wind effects. These effects 

are difficult to model and describe parametrically because of the sensitivity of the wind pressure at the 

vent locations to wind speed and direction, surrounding terrain and buildings, and fine details of the 

building architecture. Thus, we have not included wind effects in our analysis. One approach to 

minimizing the likelihood of adverse wind effects would be to locate the inlet (bottom) vent on the 

prevailing upwind side of the building and the outlet (top) vent on the prevailing downwind side. In 

this configuration, the prevailing winds would most likely augment, rather than opposing, the 

buoyancy-driven ventilation of H2 from the building. 

6.0 MECHANICAL VENTILATION 

Because buoyancy-driven passive ventilation has limitations, we also consider the alternative 

approach of mechanical ventilation. With reference to Fig. 1 and Eqn. 5, we calculate the required 

airflow rate, QT, through an exhaust fan placed in the top vent as a function of the leakage rate, S, and 

the concentration at the outlet, cT. IEC Standard 60079-10 [23] presents a similar formulation, which 

also accounts for airflow impediments and a temperature correction, with the result expressed in terms 

of an effective cloud volume based on the actual net ventilation rate. With our formulation, applying 

the safe limit value of cT = 1%, the indicated airflow rates are shown in Table 2. For perspective, we 

note that airflow rates for bathroom exhaust fans and whole-house ventilation systems are typically in 

the range of about 25 to 50 L/s. For leak-down times of 1 day and less, the ventilation rates listed in 

Table 2 are significantly higher than this range. Constant operation of such a fan does not seem 

practical, because of both the fan energy consumption and the effect on garage temperature in cold 

weather. Thus, a fan controller based on H2 detection would be recommended, and the reliability of 

the control system would be essential to safety. 

Table 2. Fan sizing chart for mechanical ventilation, based on cT = 1%. 

Fan Airflow

T, days T, hrs L/min L/s

0.25 6 166 276

0.5 12 82.9 138

1 24 41.5 69

2 48 20.7 35

3 72 13.8 23

7 168 5.92 9.9

29 696 1.43 2.4

Leakage Rate 

(Based on 5 kg of H2)

12
7.0 CONCLUSIONS 

A significant uncertainty in this analysis is the range of leakage rates that will occur with H2 vehicles. 

We have modeled and tabulated results for a range of leakage rates in order to indicate sensitivities to 

this variable. Further research is needed to ascertain the occurrence probabilities of various leakage 

rates. 

Our CFD model has not yet been compared with experimental results as a check on its accuracy. Our 

research plan includes this work as a future activity. 

The reverse thermocirculation effect can occur in nearly any climate because of steady-state or 

transient temperature differences between the garage and the ambient air. These effects can inhibit 

buoyancy-driven ventilation to the extent that H2 concentrations will exceed values predicted by the 

isothermal models. In a “perfect storm” scenario we crafted with respect to architecture, weather, and 

leak rate (Case 10), the thermal effect increased the expected H2 concentration from 2% to 5%. This 

is a significant risk factor, although the likelihood if its occurrence may be low, judging by the lengths 

we went to in order to identify a significant example. 

Based on our results to date, a reasonably conservative approach to sizing vents for passive 

ventilation of H2 is to use the isothermal vent sizing equation (Eqn. 6) with the values cT = 0.01, φ = 

2, and D = 0.6. However, the value φ = 2 is subject to CFD model validation and further case studies. 

Also, wind effects and outdoor temperatures higher than the garage temperature are additional risk 

factors, along with uncertainty about leakage rates 

Mechanical ventilation is an alternative approach to safety. Because of the large airflow rates 

indicated by Table 2, an H2-sensing fan controller is recommended. Research is needed to develop a 

control system that is sufficiently reliable and economical for residential use. 

8.0 ACKNOWLEDGMENTS 

This work is funded by the Hydrogen Safety, Codes and Standards program of the U.S. Department of 

Energy’s Office of Hydrogen, Fuels Cells and Infrastructure Technologies at the Hydrogen 

Technologies and Systems Center of the National Renewable Energy Laboratory (NREL), in 

collaboration with NREL’s Buildings and Thermal Systems Center. 

9.0 REFERENCES 

1. Bain, A., Barclay, J.A., Bose, T.K., Edeskuty, F.J., Fairlie, M.J., Hansel, J.G., Hay, D.R., Swain, 

M.R., Ohi, J.M. and Gregoire Padró, C.E. Sourcebook for Hydrogen Applications, 1998, TISEC, 

Inc., Montreal, Canada. 

2. ICC. International Fire Code, Section 2211.7.1.1, 2006, International Code Council, 

Washington, DC. http://www.iccsafe.org, last accessed 6/5/07. 

3. NFPA. Vehicular Fuel Systems Code, Section 9.3.3.5.4.1, 2006, National Fire Protection 

Association, Quincy, MA. http://www.nfpa.org/, last accessed 6/12/07. 

4. Swain, M.R., Filoso, P.A. and Swain, M.N. Ignition of Lean Hydrogen-Air Mixtures, Int. J. 

Hydrogen Energy, 30, No. 30, 2005, pp. 1447-1455. 

5. Swain, M.R., Filoso, P.A. and Swain, M.N. An Experimental Investigation into the Ignition of 

Leaking Hydrogen, Int. J. Hydrogen Energy, 32, No. 2, 2007, pp. 287-295. 

6. Standard SAE J2578. Surface Vehicle Recommended Practice, 2002, SAE International, 

Warrendale, PA. http://www.sae.org/servlets/index, last accessed 4/27/07. 

13
7. Cadwallader, L.C. and Herring, J.S. Safety Issues with Hydrogen as a Vehicle Fuel, Idaho 

National Engineering and Environmental Laboratory, Lockheed Martin Idaho Technologies 

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8. Swain, M.R. and Swain, M.N. Passive Ventilation Systems for the Save Use of Hydrogen, Int. J. 

Hydrogen Energy, 21, No. 10, 1996, pp. 823-835. 

9. Breitung, W., Necker, G., Kaup, B. and Veser, A. Numerical Simulation of Hydrogen Release in 

a Private Garage, Proceedings of Hypothesis IV, 9-14 September 2001, Strahlsund, Germany, p. 

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10. Swain, M.R., Filoso, P., Grilliot, E.S. and Swain, M.N. Hydrogen Leakage into Simple 

Geometric Enclosures, Int. J. Hydrogen Energy, 28, No. 2, 2003, pp. 229-248. 

11. Swain, M.R. Fuel Leak Simulation. Proceedings of the 2001 U.S. DOE Hydrogen Program 

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http://www.nrel.gov/docs/fy01osti/30535.pdf. 

12. Papanikolaou, E.A. and Venetsanos, A.G. CFD Modeling for Helium Releases in a Private 

Garage without Forced Ventilation, Proceedings of the Fist ICHS, Pisa, Italy, 8-10 September 

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13. Emswiler, J.E. The Neutral Zone in Ventilation. Trans. Amer. Soc. Heat. Vent. Engrs., 32, 1926, 

pp. 59-74. 

14. Brown, W.G. and Solvason, K.R. Natural Convection through Rectangular Openings in 

Partitions—1. Int. J. Heat Mass Transfer, Vol. 5, 1962, pp. 859-868. 

15. Fitzgerald, S.D. and Woods, A.W. Natural Ventilation of a Room with Vents at Multiple Levels. 

Building and Environment, Vol. 39, 2004, pp. 505-521. 

16. ASHRAE. Handbook—Fundamentals, Chapter 27, Eqn. 28. American Society of Heating, 

Refrigerating, and Air-Conditioning Engineers, Inc., 2005, Atlanta, GA. 

17. Building America homepage, http://www.eere.energy.gov/buildings/building_america/, last 

accessed 4/26/07. 

18. Fluent, Inc., Lebanon, NH, http://www.fluent.com/, last accessed 4/26/07. 

19. Idelchik, I.E. Handbook of Hydraulic Resistance, 2nd Ed, Diagram 4-18. Hemisphere Publishing 

Corp., 1986, New York, NY. 

20. Building Science Consulting. Discussion of the Use of Transfer Grilles to Facilitate Return Air 

Flow in Central Return Systems, 

http://www.buildingscienceconsulting.com/resources/mechanical/hvac/transfer_grills.htm, last 

accessed 4/26/07. 

21. American Lung Association. American Lung Association Health House Builder Guidelines, 

Effective 10/1/04, The American Lung Association, 2004, New York, NY. 

22. Lindeburg, M. Mechanical Engineering Review Manual, 7th ed., Professional Publications, Inc., 

1984, Belmont, CA, pp. 6-21. 

23. IEC. International Standard 60079-10: Electrical apparatus for explosive gas atmospheres, 

Annex B, Sec. B.4.2.2, 2002, International Electrotechnical Commission, Geneva, Switzerland. 

http://www.iec.ch/, last accessed 6/19/07. 

14
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Analysis of Buoyancy-Driven Ventilation of Hydrogen from Buildings: 

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14. ABSTRACT (Maximum 200 Words)

When hydrogen gas is used or stored within a building, as with a hydrogen-powered vehicle parked in a residential 

garage, any leakage of unignited H2 will mix with indoor air and may form a flammable mixture. One approach to 

safety engineering relies on buoyancy-driven, passive ventilation of H2 from the building through vents to the outside.

15. SUBJECT TERMS

hydrogen; buoyancy; passive ventilation 

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