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Annals of Occupational Hygiene Advance Access originally published online on July 20, 2006
Annals of Occupational Hygiene 2007 51(1):35-43; doi:10.1093/annhyg/mel049
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© The Author 2006. Published by Oxford University Press on behalf of the British Occupational Hygiene Society

Local Ventilation Solution for Large, Warm Emission Sources

ILPO KULMALA1, PASI HYNYNEN2, IRMA WELLING2,* and ARTO SÄÄMÄNEN1

1 VTT Industrial Systems P.O. Box 13001, FIN-33101 Tampere, Finland
2 Finnish Institute of Occupational Health Laserkatu 6, FIN-53850 Lappeenranta, Finland

*Author to whom correspondence should be addressed. Tel: +358 304 743201; fax: +358 304 743230; e-mail: irma.welling{at}ttl.fi


   ABSTRACT
TOP
 ABSTRACT
INTRODUCTION
 MATERIALS AND METHODS
 RESULTS AND DISCUSSIONS
 CONCLUSIONS
 APPENDIX
 ACKNOWLEDGEMENTS
 REFERENCES
 
In a foundry casting line, contaminants are released from a large area. Casting fumes include both volatile and particulate compounds. The volatile fraction contains hydrocarbons, whereas the particulate fraction mostly comprises a mixture of vaporized metal fumes. Casting fumes lower the air quality in foundries. The design of local ventilation for the casting area is a challenging task, because of the large casting area and convection plumes from warm moulds. A local ventilation solution for the mould casting area was designed and dimensioned with the aid of computational fluid dynamic (CFD) calculations. According to the calculations, the most efficient solution was a push–pull ventilation system. The prototype of the push–pull system was built and tested in actual operation at the foundry. The push flow was generated by a free plane jet that blew across the 10 m wide casting area towards an exhaust hood on the opposite side of the casting lines. The capture efficiency of the prototype was determined by the tracer gas method. The measured capture efficiencies with push jet varied between 40 and 80%, depending on the distance between the source and the exhaust. With the aid of the push flow, the average capture efficiency was increased from 40 (without jet) to 60%.

Keywords: CFD • push-pull ventilation • casting • local exhaust ventilation


   INTRODUCTION
TOP
 ABSTRACT
 INTRODUCTION
MATERIALS AND METHODS
 RESULTS AND DISCUSSIONS
 CONCLUSIONS
 APPENDIX
 ACKNOWLEDGEMENTS
 REFERENCES
 
In foundries, casting fumes spread by convection plumes include both volatile and particulate compounds. In casting, the molten hot metal causes evaporation and reaction of some organic compounds from the cast mould. Toxic compounds, such as carbon monoxide (CO) and polycyclic aromatic hydrocarbons (PAHs), can be formed. These toxic compounds and vaporized metal fumes spread by the convection plumes present a risk to the employees' health. The fumes can induce upper respiratory tract and eye irritation, and also asthma.

In order to reduce the workers' exposure, the airborne contaminants should be removed with an effective local ventilation system. However, pollutants having buoyancy and released from a wide area are the challenges for a ventilation system. It is well known that the control distance of an exhaust hood is very limited. The use of a simple overhead canopy hood is also limited owing to overhead crane access requirements. One possible solution for controlling such large sources is a push–pull ventilation system, where a jet of air is blown from one side of the source area and air is sucked by a hood on the opposite side. Along the path of the air, jet pollutants are induced by the air jet and carried into the exhaust hood. Basic design values for push–pull ventilation systems have been given by the American Conference of Governmental Industrial Hygienists (ACGIH, 1995). A push–pull ventilation system has been designed for: foundry/welding (Komine et al., 1997), solder fumes in the electronics industry (Cherrie et al., 1997; Watson et al., 2001) and open tanks (Robinson and Ingham, 1995; Woods and McKarns, 1995; Marzal et al., 2002a,b; 2003a,b). Robinson and Ingham (1996) compared existing design instructions and derived recommendations for push–pull systems where the supply air forms a 2D wall jet.

Computational fluid dynamics (CFD) has been applied to push–pull ventilation systems (Flynn et al., 1995; Robinson and Ingham, 1996; Rota et al., 2001). CFD solutions provide good possibilities to compare different ventilation solutions. The performance of various ventilation systems as a function of different geometric configurations and operating parameters can be examined. The results of such studies improve the understanding of various ventilation systems and inform the design values for the push jet and exhaust flow rates.

While the present guidelines are useful for wall jet systems, they may not be accurate for situations where the supply jet forms a free jet. In addition, all previous studies have dealt with smaller dimensions. Moreover, in foundries the contaminant source is buoyant, making the flow situation and dimensioning more complex.

In this study, CFD was used in preliminary studies to assist in the selection of an optimal local exhaust solution for a casting area in a foundry. On the basis of the results, a push–pull solution was selected to be trialled. A prototype system comprising a push air jet and an exhaust hood was designed and constructed for installation in a foundry. The developed model was then validated by experimental measurements.


   MATERIALS AND METHODS
TOP
 ABSTRACT
 INTRODUCTION
 MATERIALS AND METHODS
RESULTS AND DISCUSSIONS
 CONCLUSIONS
 APPENDIX
 ACKNOWLEDGEMENTS
 REFERENCES
 
Description of the work, casting area and general ventilation
The study was performed in a foundry with four casting lines, each 28 m long. The width of the casting area was 10 m. In the casting, the moulds were filled with molten metal. After the mould filling the castings were allowed to cool on conveyor lines. Emissions from moulds to workplace air occurred during the cooling period, for ~4 h. PAHs can be formed during the incomplete combustion process of casting sand and binders. The main components of PAHs in the particle phase were fluoranthene, phenanthrene, pyrene and anthracene and in the gas phase the main components were naphthalene, phenanthrene, anthracene, fluorene and pyrene. PAHs are regarded as carcinogenic substances with lungs and skin the most important target organs. The workers moved on the casting area mainly during the filling of moulds.

Casting took ~15 min. Workers stayed in the casting area only during the filling of moulds. Main emissions from moulds happened during the cooling period of 4 h.

The exhaust ventilation of the foundry was arranged with roof ventilators and furnace exhausts, with a total exhaust flow rate of 76 m3 s–1. The hall was 10 m high, and allowed air contaminants to stratify with rising convection plumes to the upper part of the hall, where the air was exhausted. There were two large doors near the casting area, which were opened when raw materials were transferred inside. This caused large cold air currents to enter the casting area during heating periods.

Local ventilation solutions
In order to compare possible alternatives, CFD modelling using boundary conditions described in the next section was perfomed to study the performance of the various solutions. Opinions of the users were taken into account to make the solution workable in practice also and to make sure that it hindered working as little as possible. Four local ventilation solutions were simulated and modelled (Fig. 1):

  1. Downdraft suction between casting lines.
  2. Jet enhanced exhaust [Aaberg hood; Saunders and Fletcher (1993), Kulmala (2000) and Wen and Ingham (2000)] on the wall.
  3. Canopy hoods assisted with vertical jets.
  4. Horizontal push–pull system.
According to CFD calculations the push–pull ventilation system was the most effective. The target capture efficiency of 90% was achieved with the lowest exhaust air flow rate (1.6 m3 s–1 m–1) by the push–pull ventilation system (Fig. 2). The second most effective was the downdraft suction between casting lines with exhaust air flow rate of 1.8 m3 s–1 m–1. By using the jet enhanced exhaust (Aaberg hood) on the wall, casting fumes from the two nearest casting lines could only be removed. The highest exhaust flow rate (4.2 m3 s–1 m–1) was required for the canopy hoods assisted with vertical jets. After discussions the push–pull ventilation system was also accepted by the workers and the pilot push–pull system was dimensioned and constructed for installation. The downdraft suction solution did not fulfil the fire safety requirements owing to flying melt iron pieces.


Figure 1
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  		Fig. 1 Alternative local ventilation solutions. (A) Downdraft suction between casting lines, (B) Jet enhanced exhaust, (C) Canopy hoods assisted with vertical jets and (D) Horizontal push–pull system.

 


Figure 2
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  		Fig. 2 Capture efficiency of the different local ventilation solutions as a function of exhaust flow rate. The calculated exhaust flow rates of canopy hoods for efficient operation were about twice that of the other solutions.

 
CFD modelling
Numerical simulations were used to determine the air flow fields and to evaluate efficiencies of the alternative solutions. The numerical studies were made with FLUENT version 4.5 code, assuming 3D, steady-state and non-isothermal flow. The CFD code solves the conservation equations of mass, momentum, energy and turbulence quantities using a method based on a control-volume. For turbulence the standard k-{varepsilon} model was applied.

Because of computational resource limitations, only one slice of the casting line was modelled using a calculation grid of 48 x 20 x 98 cells. Further grid refinements were not possible so that the sensitivity of model results to the grid size was estimated by comparing the dominant flow features, such as jet and buoyant flows, with analytical solutions. The computation domain and boundary conditions used in the simulations are shown in Fig. 3. In modelling of local exhausts and air jets, constant velocity was assumed at the exhaust and supply openings.


Figure 3
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  		Fig. 3 CFD calculation domain and boundary conditions.

 
Modelling of convection plume
In the simulations of convection plumes a constant convective heat release rate from the moulds was assumed. This heat release rate was calculated from the equation:

Formula 1(1)
where h is the convection heat transfer coefficient, A is the area of the mould surface, Ts is the surface temperature on the mould and Ta is the ambient air temperature.

The heat transfer coefficient (h) was obtained from the measured mean surface and ambient temperatures using an empirical expression for a vertical plate (Churchill and Chu, 1975):

Formula 2(2)
where NuL is the Nusselt number, RaL is the Rayleigh number and Pr is the Prandtl number. These numbers and volumetric thermal expansion coefficient ß are defined by equations (3)–(6):

Formula 3(3)

Formula 4(4)

Formula 5(5)

Formula 6(6)
where Lm is the height of the mould, kf is the thermal conductivity of the air at film temperature, g is the acceleration of gravity and {nu} is the kinematic viscosity and {alpha} thermal diffusivity of air. On the basis of measured surface temperatures, the average convective heat release rate was estimated to be 1.5 kW for each mould. The buoyant flow calculations were made for the cooling moulds using the estimated convective heat release rates. These values are rough averages of the different size moulds and they do not take into consideration the time-dependent behaviour during cooling. However, as the upward velocity of the buoyant plume is proportional to Formula 6, it was assumed that the effect of errors in the heat flow rate estimations on the predicted flow field is not very significant for practical flow calculations.

The buoyant flow will interact with the jet flow creating a complex 3D flow field. In order to adequately predict the resulting flow, it is necessary that the buoyant flow is modelled satisfactorily. The ability of the k-{varepsilon} model to predict buoyant flows was, thus, studied by calculating buoyant flows separately and comparing the results with existing data. After this, the whole situation was modelled taking into consideration buoyant flows, the jet and the exhaust.

Modelling of the air jet
In push–pull systems, the supply jet has a crucial effect on the performance of the solution. In the pilot push–pull system the supply jet was issued from a narrow slot with a large aspect ratio so that it could be approximated by a 2D free plane jet. The dimensioning of such a system differs somewhat from that of conventional push–pull systems where the jet is a wall jet attached to a surface.

The jet entrains air as it is issued at a relatively high velocity from a narrow slot. As the jet issues, a shear layer is formed on both sides of the jet, which tends to slow down the velocity and entrain ambient air, increasing the flow volume rate. The jet spreads linearly with downstream distance and the characteristic jet width is given by

Formula 7(7)
where y0.5 denotes the location where the jet velocity is half of its maximum (centreline) value and z the distance from the jet opening (Fig. 4).


Figure 4
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  		Fig. 4 Predicted and experimental jet centerline velocity decay.

 
The jet velocity decreases with the downstream distance. The mean centreline velocity WCL can be calculated by (Chen and Rodi, 1980):

Formula 8(8)
where I is the momentum of the jet per length given by

Formula 9(9)
where W0 is the jet velocity at the jet opening. The required control velocity at the contaminant source depends on the characteristics of the contaminant source and disturbing air currents. In the case of large sources the control velocity varies since close to the jet the velocity is greater than near the exhaust. On the basis of necessary control velocities, Robinson and Ingham (1996) recommend that for locations with typical cross-draughts the supply jet momentum should be

Formula 10(10)
where z is the distance between jet nozzle and source. These correspond to capture velocities of 0.9–1.2 m s–1.

The jet of the pilot system was too narrow to be simulated in detail, because this would have resulted in excessive number of computational cells with undesirable large aspect ratios. Therefore the jet was approximated with a wider jet with the same momentum, since this is the most important parameter for the jet (Robinson and Ingham, 1996). The velocity at the jet exit WL was thus calculated by

Formula 11(11)
where b0 is the nozzle width of the actual jet (0.0034 m) and b1 the width of the jet in the model. In the simulations the jet was modelled assuming a jet width of 0.1 m. The predicted centreline velocity of the jet is compared with experimental values in Fig. 4. It can be seen that although the calculated velocities are underpredicted near the jet exit, the agreement is fairly good in regions where the contaminant sources located.

As a result of entrained air, the jet flow rate increases with distance. For undisturbed conditions the jet flow rate at the exhaust hood can be calculated by (Zhihov, 2001)

Formula 12(12)
where L is the distance between the jet nozzle and exhaust hood. This is the flow that must be exhausted by the pull hood. Inserting the values used in the foundry (W0 = 20 m s–1, b0 = 0.0034 m, distance L = 8.2 m) we get air flow rates between 1.43 and 1.77 m3 s–1 per unit width. The final exhaust flow rate was fine tuned by calculating the capture efficiency of the push–pull system using different pull flow rates.

On the basis of the simulations a pilot push–pull ventilation system was designed and constructed. The system consisted of a 5 m wide and 3.4 mm high supply plane jet blowing across the casting area towards a 6 m wide and 0.8 m high exhaust hood on the opposite side of the casting lines. To maintain uniform velocity at the exhaust the air was sucked through perforated plate with 10% open area. The predicted exhaust flow rate was 1.6 m3 s–1 per unit width for efficient operation under undisturbed conditions. This is the flow rate estimation, which was used for dimensioning and selecting the exhaust fan.

Measurements and data acquisition
Environmental measurements were performed before applying the local ventilation solution and for validation of the pilot local ventilation solution. Air contaminant, temperature and air velocity measurements were performed to characterize the source area and environment for the CFD calculations. Air movements were visualized by smoke.

The air temperature was measured at one point between casting lines 2 and 3 at four heights (0.4, 1.4, 2.4 and 3.4 m) with a real-time system. In addition, the air temperature in the supply air jet was followed after the installation of the local ventilation system. The surface temperatures of the moulds were also recorded. Air velocity was measured above the moulds with anemometers (Kaijo Denki 3D ultrasound, Japan, Airflow Developments rotating vane anemometer, UK).

The performance of the local ventilation system was investigated using the tracer gas method. Tracer gas (sulphur hexafluoride SF6) was released on the hot moulds at four casting lines to simulate the dispersion of the contaminants by warm plumes. The concentration of the tracer was measured with real-time gas analysers (Binos, USA, and Brüel & Kjaer 1302 + 1303, Denmark) from the local exhaust duct. The background concentration was monitored outside the area of influence of the local ventilation. Tracer gas was also released directly into the local exhaust duct to achieve the reference concentration. The reference concentration corresponds to the concentration in the exhaust duct with 100% capture efficiency. The background concentration was subtracted from the measured concentrations as shown in equation (13),

Formula 13(13)
where Formula 13 is the measured concentration in the local exhaust, Ca1 and Ca2 are background concentrations in the exhaust before and after the tracer release and Ca3 is the background concentration at the suction area of air jet. The reference concentration (Cref) was calculated similarly.

The capture efficiency ({eta}) was calculated as follows:

Formula 14(14)


   RESULTS AND DISCUSSIONS
TOP
 ABSTRACT
 INTRODUCTION
 MATERIALS AND METHODS
 RESULTS AND DISCUSSIONS
CONCLUSIONS
 APPENDIX
 ACKNOWLEDGEMENTS
 REFERENCES
 
Modelling results
The calculated pure plume velocities are shown in Fig. 5. Previous plume flow studies have shown that for an axisymmetric plume in a uniform environment the mean vertical centreline velocity UCL varies in the decay region as

Formula 15(15)
where g is the gravitational acceleration, {varphi}CONV is the convective heat release rate and x is the vertical distance above the virtual origin (xo) of the plume.


Figure 5
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  		Fig. 5 Comparison of calculated and empirical mean centreline velocities along the axis of plume.

 
Although there is a consensus about the general form of the plume, there is some difference among results from different researchers. Shabbir and George (1994) obtained value K = 3.4 in their extensive plume measurements, and this was used in the comparisons. In addition, in the case of extended sources there are variations in the location of the virtual origin.

It can be seen that above the heat source the fluid first accelerates and then starts to decelerate. In the decay region the predicted velocity is somewhat overestimated with the standard k-{varepsilon} model but the overall behaviour is predicted quite well. In a previous study (Welling, 2000), the velocities measured above an extended source (vertical cylinder with a diameter of 0.32 m), the velocities at x/D = 1.6, were already high, whereas in the predictions the velocities at similar distances were still accelerating (D mould = 0.5 m).

The predicted mean velocity and contaminant contours for this solution are shown in Figs 6 and 7. It can be seen from Fig. 6 that near the nozzle exit the initial isothermal jet flow is horizontal. As the jet flows over the hot moulds, the entrained warm air causes the jet to bend upwards at the same time as its velocity decreases downstream. Near the opposite wall the jet flow is removed by the exhaust hood. The predicted time-averaged concentration fields in Fig. 7 show clearly the transport of contaminants from the moulds to the exhaust hood. Because of the high velocities close to the jet nozzle, the contaminant flow from the two farthest moulding lines from the exhaust hood is abruptly bent towards the exhaust hood. The contaminants are dispersed in the jet flow owing to turbulent diffusion. The entrainment and mixing process continues towards the exhaust as shown in Fig. 7.


Figure 6
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  		Fig. 6 Predicted constant velocity contours in the vertical symmetry plane. Velocities are expressed in m s–1. The location of the temperature measurement points are also shown.

 


Figure 7
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  		Fig. 7 Predicted contaminant contours in the vertical symmetry plane. The figures depict relative concentration.

 
The predictions describe the mean flow behaviour only and do not solve the very complex time-dependent eddy currents in the shear layers of both the buoyant and jet flow. These turbulent eddies may be responsible for the enhanced mixing of contaminants.

Smoke visualizations
The smoke visualization technique helped to understand the behaviour of the push flow (Fig. 8). With smoke it was easy to see whether the push flow was bent too much, causing the escape of contaminants from the exhaust hood. The smoke tests showed also that the overall velocity patterns are similar to those predicted.


Figure 8
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  		Fig. 8 Smoke was released to push air jet to study the behaviour of the push flow.

 
Optimization of the push air jet
In the optimization of the supply air, the flow rate of the local exhaust was 8.0 m3 s–1. The capture efficiency was determined by four air flow rates of the push air jet (0.2, 0.3, 0.4 and 0.5 m3 s–1) and by releasing tracer gas from casting line 3. The results (Fig. 9) indicated that the highest mean capture efficiency of all lines was achieved by air flow rate of 0.35 m3 s–1.


Figure 9
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  		Fig. 9 The optimization of push air jet flow rate.

 
Capture efficiencies
The predicted capture efficiencies were clearly higher than measured. It is difficult to say whether the discrepancies are due to the underprediction of turbulent dispersion or due to the failure to take into consideration all the important disturbances at the freestream boundaries. In reality, the ambient air currents in the industrial hall may reduce the efficiency.

Capture efficiencies were measured in two different cases:

  • Case 1: local exhaust 8.9–9.1 m3 s–1 and supply air jet 0.35 m3 s–1.
  • Case 2: local exhaust 8.9–9.1 m3 s–1 and supply air jet 0 m3 s–1.

The capture efficiency increased in both cases while the tracer gas release point was moved from the farthest line to the nearest line to the local exhaust. The measured capture efficiencies with push jet varied between 40 and 80%, depending on the distance between the source and the exhaust. It was shown that significantly better capture efficiencies were achieved with lines 1–2 using a supply air jet than without it. On average, the push flow increased the average capture efficiency from 40 to 60% showing the positive effect of supply jet on the performance of the ventilation system.


   CONCLUSIONS
TOP
 ABSTRACT
 INTRODUCTION
 MATERIALS AND METHODS
 RESULTS AND DISCUSSIONS
 CONCLUSIONS
APPENDIX
 ACKNOWLEDGEMENTS
 REFERENCES
 
The design of an effective local ventilation system for large contaminant sources with buoyancy is a challenging task. CFD simulations were found to be a useful tool for studying different alternatives and for designing the local ventilation solutions. However, in reality there are often complex, unsteady air flow patterns, which are very difficult to take into consideration in the simulations. Because the flow field in shear layers of jet flow is very complex and filled with vortex structures, modelling is very difficult.

According to CFD calculations the push–pull ventilation system was the most effective and the target capture efficiency of 90% was achieved with the lowest exhaust air flow rate by the push–pull ventilation system. The second effect was the downdraft suction between casting lines. By using the jet enhanced exhaust (Aaberg hood) on the wall, casting fumes from the two nearest casting lines could be removed. The highest exhaust flow rate was required for the canopy hoods assisted with vertical jets. In practice the measured capture efficiencies remained lower than 90%.

These studies showed that the effective control range can be increased significantly by using assisting jets. However, for optimum performance there should be a correct balance between push jet and exhaust. The push air jet flow increased the average capture efficiency from 40 to 60%. The biggest increases were measured with lines 1 and 2. The push jet air flow creates an area in which air movements are controlled and velocities are higher than the convective flow velocities and ambient disturbances. The common worry of workers was that the air push plenum was interfering with work during filling of the moulds. To make the air push plenum more user-friendly, it ought to be divided into 2–3 m long intervals with 1 m distance between.

The highest capture efficiencies were measured from casting line 4, which was nearest to the exhaust. The most useful casting arrangement is to concentrate castings as much as possible on lines nearest to the exhaust hood. Although the average capture efficiency of 60% may be deemed satisfactory, taking into consideration the large size of the source, it was less than the targeted value of 90%. It may be concluded that the predictions underestimate the dispersion of contaminants, perhaps because of the inability to model unsteady turbulent diffusion properly. Because the flow field in shear layers of jet flow is very complex and filled with vortex structures, modelling is very difficult. Therefore, the calculation results should be considered cautiously and for design purposes a safety factor should perhaps be used.

For the studied case the design value for the exhaust flow rate was 1.6 m3 s–1 m–1 of exhaust hood. About 10% perforated material is suggested for the exhaust. The design value of the push air flow rate was 0.07 m3 s–1 m–1 of the slot. The correct height of the slot was 3.4 mm and the air velocity in the slot was 20 m s–1.

We conclude that when push air jet and exhaust are combined in a correctly balanced ratio, controlled air movement can be obtained over much greater distances than is possible with conventional systems.


   APPENDIX
TOP
 ABSTRACT
 INTRODUCTION
 MATERIALS AND METHODS
 RESULTS AND DISCUSSIONS
 CONCLUSIONS
 APPENDIX
ACKNOWLEDGEMENTS
 REFERENCES
 


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   ACKNOWLEDGEMENTS
TOP
 ABSTRACT
 INTRODUCTION
 MATERIALS AND METHODS
 RESULTS AND DISCUSSIONS
 CONCLUSIONS
 APPENDIX
 ACKNOWLEDGEMENTS
REFERENCES
 
We are grateful to Sulzer Pumps Finland and industrial hygienist Juhani Reponen from Ahlström Karhulan Palvelut for help with the evaluation of different solutions, installation of the push–pull ventilation system and allowing us to collect data at their workplace. The technical assistance of occupational technicians Pertti Närhi and Timo Nurkka and laboratory engineer Timo Mielo, Finnish Institute of Occupational Health, is gratefully acknowledged. The work was funded by the Finnish Work Environment Fund.

Received December 30, 2005; in final form June 15, 2006


   REFERENCES
TOP
 ABSTRACT
 INTRODUCTION
 MATERIALS AND METHODS
 RESULTS AND DISCUSSIONS
 CONCLUSIONS
 APPENDIX
 ACKNOWLEDGEMENTS
 REFERENCES
 

ACGIH. (2001) Industrial ventilation, a manual of recommended practice. 24th edn (American Conference of Governmental Industrial Hygienists, Cincinnati, OH).

Chen CJ and Rodi W. (1980) Vertical turbulent buoyant jets—a review of experimental data. (Pergamon Press, Oxford).

Cherrie JW, Hamid EA, O'Hara V. (1997) A push–pull ventilation system for use in hand soldering. Proceedings of Ventilation 97' Vol. 12: pp. 658–65.

Churchill SW and Chu HHS. (1975) Correlating equations for laminar and turbulent free convection from a vertical plate. Int J Heat Mass Transfer 18:1323–9.[CrossRef]

Flynn MR, Kwangseog A, Miller CT. (1995) Three-dimensional finite-element simulation of a turbulent push–pull ventilation system. Ann Occup Hyg 39:573–89.[Abstract/Free Full Text]

Komine H, Tsuji K, Mori Y. (1997) Push and pull ventilation system application for foundry/welding. Proceedings of Ventilation 97' Vol. 12: pp. 1121–30.

Kulmala I. (2000) Experimental validation of potential and turbulent flow models for a two-dimensional jet enhanced exhaust hood. Am Ind Hyg Assoc J 61:183–91.

Marzal F, Gonzalez E, Minana A, et al. (2002a) Influence of push element geometry on the capture efficiency of push–pull ventilation systems in surface treatment tanks. Ann Occup Hyg 46:383–93.[Abstract/Free Full Text]

Marzal F, Gonzalez E, Minana A, et al. (2002b) Determination and interpretation of total and transversal linear efficiencies in push–pull ventilation systems for open surface tanks. Ann Occup Hyg 46:629–35.[Abstract/Free Full Text]

Marzal F, Gonzalez E, Minana A, et al. (2003a) Methodologies for determining capture efficiencies in surface treatment tanks. Am Ind Hyg Assoc J 64:604–8.

Marzal F, Gonzalez E, Minana A, et al. (2003b) Visualization of airflows in push–pull ventilation systems applied to surface treatment tanks. Am Ind Hyg Assoc J 64:455–60.

Robinson M and Ingham DB. (1996a) Recommendations for the design of push–pull ventilation systems for open surface tanks. Ann Occup Hyg 40:693–704.[Abstract/Free Full Text]

Robinson M and Ingham DB. (1996b) Numerical modelling of the flow patterns induced by a push–pull venilation system. Ann Occup Hyg 40:293–310.[Abstract/Free Full Text]

Rota R, Nano G, Canossa L. (2001) Design guidelines for push–pull venilation systems through computational fluid dynamics modeling. Am Ind Hyg Assoc J 62:141–8.

Saunders CJ and Fletcher B. (1993) Jet enhanced local exhaust ventilation. Ann Occup Hyg 37:15–24.[Abstract/Free Full Text]

Shabbir A and George WK. (1994) Experiments on a round turbulent buoyant plume. J Fluid Mech 275:1–32.[CrossRef]

Watson SI, Cain JR, Cowie H, et al. (2001) Development of a push–pull ventilation system to control solder fume. Ann Occup Hyg 45:669–79.[Abstract/Free Full Text]

Welling I. (2000) Experimental study of natural-convection plumes from a heated horizontal square plate and a vertical cylinder. Exp Heat Transfer 13:7–19.

Wen X and Ingham DB. (2000) Theoretical and numerical predictions of two-dimensional Aaberg slot exhaust hoods. Ann Occup Hyg 44:375–90.[Abstract/Free Full Text]

Woods JN and McKarns JS. (1995) Evaluation of capture efficiencies of large push–pull ventilation systems with both visual and tracer techniques. Am Ind Hyg Assoc J 56:1208–14.

Zhihov A. (2001) Air jets. In Goodfellow H and Tähti E (Eds.). Industrial ventilation. design guidebook(Academic Press, London) pp. 446–512.


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