Dimension values of the full-scale model shown in \nFigure 6\n.
\r\n\t[2] J. V. Moloney, A. C. Newell. Nonlinear Optics. Westview Press, Oxford, 2004.
\r\n\t[3] M. Kauranen, A. V. Zayats. Nonlinear Plasmonics. Nature Photonics, vol. 6, 2012, pp. 737-748.
\r\n\t[4] P. Dombi, Z. Pápa, J. Vogelsang et al. Strong-field nano-optics. Reviews of Modern Physics, vol. 92, 2020, pp. 025003-1 – 025003-66.
\r\n\t[5] N. C. Panoiu, W. E. I. Sha, D.Y. Lei, G.-C. Li. Nonlinear optics in plasmonic nanostructures. Journal of Optics, 20, 2018, pp. 1-36.
\r\n\t[6] A. Krasnok, A. Alu. Active nanophotonics. Proceedings of IEEE, vol. 108, 2020, pp. 628-654.
\r\n\t[7] M. Lapine, I.V. Shadrivov, Yu. S. Kivshar. Colloquium: Nonlinear metamaterials. Reviews of Modern Physics, vol. 86, 2014, pp. 1093-1123.
\r\n\t[8] Iam Choon Khoo. Nonlinear optics, active plasmonics and metamaterials with liquid crystals. Progress in Quantum Electronics, vol. 38, 2014, pp. 77- 117.
\r\n\t
Pump intake is the part of a pump that draws fluid from the reservoir called the sump as a result of pressure difference generated by the impeller. In most cases, pumped fluid enters the intake in a swirling motion due to geometric features of the sump [1]. Inappropriate sump design such as abrupt changes in sump boundaries, narrow clearance under the pump inlet and asymmetric orientation of the approach channel to the sump will lead to the formation of swirls and vortices [2]. Strong vortices may cause damages to the pump impeller by channelling air to the impeller surface and initiate adverse effects such as cavitation and vibration [3]. On the other hand, excessive swirls in the intake flow can impose imbalance loading to the impeller and even bring resistance to the impeller rotation by introducing swirl rotation in the opposite direction [4]. Due to site condition and operational restrictions, optimal sump design may not be achieved, and therefore local flow correction devices are used as remedial measures.
\nThese devices which are commonly known as anti-vortex device (AVD) come in different shapes and sizes, depending on its application. The conceptual design of AVD is outlined in ANSI/HI 9.8-2018 [5] standard which is a guideline to assist engineers and designers in optimal intake sump design. Among the AVD types employed in real applications are floor splitter [6], floor cone [7] and corner fillet [8]. These AVD types serve the purpose of eliminating submerged vortices formed at the sump floor. Floor splitters are the most widely used AVD type due to its effectiveness in eliminating vortices and reducing vorticity in the pump intake flow. There are two versions of floor splitter, namely the prism and the plate types. The use of plate type floor splitter is favourable in many applications due to its fabrication friendly-feature and economic design [9]. However, there are a limited number of articles in the literature which discuss the features of floor splitter plate in detail. In this chapter, the characteristics of swirl angle reduction of a floor splitter plate installed in pump sump are studied.
\nThe study was carried out by both experimental and numerical approaches. A single intake pump sump model, as shown in \nFigure 1\n, was utilized for the study in which the sample of a floor splitter was installed beneath an intake suction pipe in the sump model. The layout of the sump model test section and the dimensions of the floor splitter installed is illustrated in \nFigure 2(a)\n and \n(b)\n, respectively.
\nThe experimental rig.
Main dimensions of the sump model and splitter.
The main objective of the study is to evaluate the swirling motion in the intake pipe and associated with submerged vortex without and with the installation of floor splitter plate. Initially, the experiment was conducted without the installation of floor splitter plate to capture the initial conditions of the setup. The measurement of the intensity of swirl in the intake pipe was performed according to the procedure described in ANSI/HI 9.8-2018 standard for pump sump model test. The parameter used to quantify the measurement data is the swirl angle θ which is defined in the following equation:
\nwhere d is the inner diameter of the intake pipe, n is the revolution count of the measurement instrument called the swirl metre and a is the average axial velocity at the location of the swirl metre. The swirl metre consists of a shaft with four straight blades used to capture the swirling motion in the intake pipe, and the revolution count of the swirl metre blade is measured using a tachometer. \nFigure 3\n shows typical swirl metre installation according to ANSI/HI 9.8-2018 standard. Basically, θ is the convention for describing the ratio between the axial velocity and the tangential velocity of the intake flow which characterizes the intensity of the swirling motion in the fluid. The acceptance criteria according to ANSI/HI 9.8-2018 is that the swirl angle must be lower than 5° to prevent excessive swirl in the intake flow.
\nSwirl metre installation according to ANSI/HI 9.8-2018 standard.
In order to generate the submerged vortex, the clearance under the pipe was set to 0.3 times the diameter of the inlet D and two types of flow conditioners were installed: a sloped floor with an inclination angle of 30° and a sloped wall with the same inclination angle. These flow conditioners were installed at a distance of about 5D from the centre of the intake pipe as shown in \nFigure 4(a)\n and \n(b)\n, respectively. The measurement was conducted in a range of pump submergence levels which are normalized by the minimum inlet submergence Smin\n, a threshold value before the occurrence of a surface vortex. Smin\n is calculated by the following equation:
\nFalse floor and false wall arrangements.
\nFrin\n is the Froude number at the pipe inlet and is given by:
\nwhere νin\n is the flow velocity at the inlet and g is the gravitational acceleration. The range of the dimensionless parameter S/Smin\n was set between 0.8 and 1.2.
\nThe numerical approach part of the study is set for the simulation of the flow in a full-scale pump sump. As the construction cost for a full-scale pump sump cannot be afforded, a computational fluid dynamics (CFD) simulation was employed as a replacement. The numerical model was validated with experimental data and incorporated with a combined flow conditioner that consists of inclined floor and inclined wall as the ones used in the experiment and built at a scale of 9:1. The flow rate of the pump was set to 2170 l/s, and the pump submergence took the value of Smin\n which is, after the calculation by using Eq. (2), equals to 2678 mm. The mesh structure and the dimensions of the full-scale pump sump are illustrated in \nFigures 5\n and \n6\n, respectively, while the values of the model dimensions are listed in \nTable 1\n.
\nNumerical model of the full-scale pump sump; (a) the computational domain, (b) model without floor splitter plate, (c) model with floor splitter plate.
Dimensions of the full-scale model.
Parameter | \nDimension (mm) | \n
---|---|
Inlet diameter D\n | \n1275 | \n
Pipe diameter d\n | \n850 | \n
Right side distance W1 | \n1190 | \n
Left side distance W2 | \n1360 | \n
Water entrance width W3 | \n1275 | \n
Intake pipe height H1 | \n9350 | \n
Sump height H2 | \n4250 | \n
Water entrance height H3 | \n2975 | \n
Floor length L1 | \n6375 | \n
Water entrance distance from sloped floor L2 | \n8417 | \n
Clearance C\n | \n382.5 | \n
Dimension values of the full-scale model shown in \nFigure 6\n.
\n\nFigures 7\n and \n8\n show the distribution of swirl angle values at different submergence ratios in the case of false floor and false wall flow conditioner, respectively. Generally, the installation of floor splitter plate has shown reduction in the swirl angle values. The parameter that can be used to characterize the reduction effect of the floor splitter plate is the swirl angle reduction factor Rθ\n which is defined as follows:
\nSwirl angle values at different submergence ratios for the false floor case.
Swirl angle values at different submergence ratios for the false wall case.
In this experiment, the average value of Rθ\n for the false floor case is 1.53, while in the false wall case, the average value of Rθ\n is 1.62. In \nFigure 7\n, the swirl angle values show a decreasing trend with increasing submergence ratio for S/Smin\n greater than 1 when installed with floor splitter plate. This is due to the fact that for S/Smin\n greater than 1, there was only a submerged vortex present in the sump. As the function of floor splitter plate is to eliminate submerged vortices, this result proved that the installation of floor splitter plate has served the purpose. When S/Smin\n is decreased below 1, the swirl angle values increase with decreasing submergence ratio. The inception of free surface vortex at S/Smin\n below 1 has caused bigger fluctuation in swirl angle as can be seen in the larger uncertainties within this region. The higher swirl angle values are contributed by the increase in approach flow velocity at lower water levels. The floor splitter vortex has shown limited swirl angle reduction effect if the submergence ratio is decreased below 1.
\nIn \nFigure 8\n, the swirl angle values show a sinusoidal trend with increasing submergence ratio for S/Smin\n greater than 1 when installed with floor splitter plate. The trend is contributed by the inception of free surface vortex at S/Smin\n greater than 1. Although the theory behind the minimum inlet submergence Smin\n is that there should be no free surface vortex formed in the sump if the submergence S is greater than Smin\n, this deviation from the theory was contributed by the use of false wall in which the flow has been prerotated at the beginning of the sump. The prerotation has therefore caused the flow to develop a free surface vortex earlier than expected. In the experiment, this situation occurred at S/Smin\n = 1.15. As the swirl angle decreases when S/Smin\n decreases below 1.15, the reduction effect of the floor splitter plate can be observed in the decreasing trend of the swirl angle values. Similar to the case of false floor, the swirl angle increases as the submergence ratio decreases due to the increasing approach flow velocity at low water levels.
\nDespite the swirl angle reduction effect of floor splitter plate, the fulfilment of the requirement of swirl angle reduction below 5° has not been achieved for most of the cases. In the case of false floor, there is no submergence ratio value at which the swirl angle has been reduced below 5°; however, for the false wall case, the reduction of swirl angle values below 5° can be seen between S/Smin\n = 1.00 and S/Smin\n = 1.05, i.e. the requirement for all submergence ratios when installed with floor splitter plate. This result shows that there is a limiting factor that prevented the swirl angle reduction below 5° and that factor lies on the design of the floor splitter as suggested by Kang et al. [9].
\nThe first part of the discussion on the result of simulation of flow in full-scale pump sump model is about the vortex elimination by the installation of floor splitter plate. The evaluation is based on the vorticity in the y-axis ωy\n due to its influence on the swirling motion of the flow. The value of ωy\n is normalized by the ratio of velocity in the suction pipe and the pipe inner diameter νd/d. \nFigure 9\n shows the cross section along x-y plane in which the evaluation of the result in the streamwise direction takes place and its corresponding results which are shown in \nFigure 10\n.
\nEvaluation area in x-y plane of the pump sump model (z = 1600 mm).
Contour plot of ωy/(νd/d) at the location of vortex in x-y plane; (a) without floor splitter plate, (b) with floor splitter plate. Dashed circular lines denote pipe diameter.
From \nFigure 10\n, it can be observed that the core of the vortex, indicated by the high-intensity region extending from the floor towards inside of the pump, has been eliminated with the installation of floor splitter plate. The vorticity in the pipe has also been reduced which can be seen from the contour colors. The velocity vectors, which appear to point diagonally to the left indicating a strong swirling flow in the pipe, have been straightened in a direction vertically upward towards the direction of suction when installed with floor splitter plate.
\nWhen observing the cross section in the spanwise direction (in the plane illustrated in \nFigure 11\n), similar results are presented. Basically the flow that enters the pump is divided into two regions, namely the right side and the left side flow, due to the geometry of the sump. The flow entrance velocity from the right and the left side of the inlet are nearly the same because of the nearly symmetrical positioning of the pump. The flow entered the pump in a spiral manner without the installation of floor splitter which resulted in vortex formation near to the left side of the pump. When installed with floor splitter plate, the flow is reorganized, and therefore the spiral motion of the flow has been reduced and hence the vortex eliminated. This situation is reflected by the discontinued vortex core shown in \nFigure 12\n with the installation of floor splitter plate.
\nEvaluation area in y-z plane of the pump sump model (x = 15,860 mm).
Contour plot of ωy/(νd/d) at the location of vortex in y-z plane; (a) without floor splitter plate, (b) with floor splitter plate.
As the main function of floor splitter plate is to eliminate vortices formed at the sump floor, an evaluation about the vorticity in the plane at the sump floor is necessary. This location is shown in \nFigure 13\n. The vortex core is indicated by the spiralling streamline under the pump inlet which can be seen in \nFigure 14\n in the case without floor splitter plate. As the floor splitter plate was installed, the path of the spiral streamline was interrupted by the plate, and therefore the formation of vortex was prevented. Due to the suction by the pump, a small vortex attached to the side of the floor splitter plate was formed which is inherited from the flow without floor splitter plate as shown in \nFigure 14(b)\n. However, this vortex constitutes a much smaller vortex core diameter (estimated to be less than 0.1D based on the scale at the x-axis of the graph) and relatively weak compared to the large vortex (estimated to be about 0.2D) which can be seen in \nFigure 14(a)\n, and therefore it can be considered as nondestructive to the pump impeller.
\nEvaluation area in z-x plane of the pump sump model (y = 10 mm).
Contour plot of ωy/(νd/d) at the floor of the sump in z-x plane; (a) without floor splitter plate, (b) with floor splitter plate.
The next part of the evaluation is about the swirl angle reduction characteristics of floor splitter plate installation. For this purpose, an evaluation plane was selected at the position comparable to the installation of swirl metre in the experimental model. The location of the plane is shown in \nFigure 15\n, and its corresponding results are displayed in in \nFigure 16\n. The flow at the swirl metre location was rotational with relatively high velocity components as indicated by the velocity vectors. The two visible vorticity regions show the divided inflow field in the pipe as explained in the previous paragraph which is considerably high in reference to the value of νd/d as shown in \nFigure 16\n. With the installation of floor splitter, the magnitude of both vorticity regions is significantly reduced and the resulting velocity vectors are also smaller in size compared to the case without floor splitter. This indicates that the spiral flow has been dissolved by the floor splitter plate into a relatively straight flow and the outcome is consistent with the experimental result presented in the previous subsection.
\nEvaluation area at the position of swirl metre in z-x plane of the pump sump model (y = 4250 mm).
Contour plot of ωy/(νd/d) at the swirl metre position in z-x plane; (a) without floor splitter plate, (b) with floor splitter plate.
To get a better understanding about the result, a 3D streamline visualization of the intake flow in the sump is illustrated for every case as comparison in \nFigure 18\n. It can be seen that the intake flow was spiral before the installation of floor splitter plate and as the floor splitter was installed, the spiral motion of the flow was dissolved and went into a relatively straight path. Quantitative values can also be extracted from the result to obtain the associated swirl angle values. The approach for the calculation of swirl angle from the simulation results is based on the principle of Eq. (1) itself where by definition the swirl angle is the angle between the velocity components of the intake flow in the axial and tangential direction. From Eq. (1), the term πdn represents the tangential velocity component, while the term v represents the axial velocity component; both are at the location of the swirl metre used in the experiment. \nFigure 17\n shows the velocity triangle diagram which shows the relationship between swirl angle and both of the velocity components in a schematic representation.
\nSwirl angle definition using velocity triangle diagram as shown in Kang et al. [9].
3D streamline plot showing the intake flow in the sump with the seeding of the flow starts at the floor of the sump; (a) without floor splitter plate, (b) with floor splitter plate.
Based on this approach in Eq. (1), the velocity components in the axial and tangential direction were derived from the simulation results. As the result was given in vorticity values, the tangential velocity component must be derived from the angular velocity which equals to half of the vorticity [10]. The vorticity of the flow at the position of the swirl metre is calculated by the integration of the vorticity in the plane and divided by the cross section to obtain the vorticity value per unit area. After getting the value of angular velocity, the following correlation is used to calculate the tangential velocity:
\nThe method to derive the value of axial velocity component from the results was based on the same principle in which the integral value of axial velocity component in the plane was extracted and divided by cross-sectional area of the pipe at the swirl metre location to get the velocity per unit area. The reason of performing integration to find the velocity values is that the swirling motion of the intake flow in the pipe constitutes a solid body rotation and the swirl angle value describes the rotation body as a whole [1], and this is where the integration of the velocity across the cross-sectional area becomes the most practical way of calculating the swirl angle in the simulation. After obtaining both velocity values, the swirl angle was then calculated using the velocity triangle diagram as shown in \nFigure 17\n.
\nBy following the described procedure, the swirl angle value for the case without floor splitter plate installation is 7.58°, while for the case with floor splitter plate, the swirl angle value is 4.09°. Although these values are based on average velocities as the simulation was conducted in a steady-state simulation and therefore are much smaller than the actual swirl angle values, it can be considered as adequate because they are used for comparison purpose and not for the determination of absolute values. Once again, the results are in agreement with the experimental data. This study complements a previous experimental investigation in which the effects of floor splitter heights have been analysed [11].
\nA study on the application of vortex control principle at pump intake was carried out by using an anti-vortex device type called the floor splitter plate. The device was installed in a pump sump model to eliminate vortices formed at the intake and reduce the swirling motion in the intake pipe as a method to improve pump efficiency in actual applications. Evaluation of the effect was conducted based on experimental and numerical approaches. The experimental part comprised swirl angle measurement which was performed according to ANSI/HI 9.8-2018 standard. To complement the results obtained in the experiment, a numerical simulation of the flow in a full-scale pump sump was conducted. The results showed that the installation of floor splitter plate has successfully eliminated the vortex formed at the sump floor and reduced the swirl angle in the intake flow. However, the reduction effect was not sufficient to achieve the criteria set in the ANSI/HI 9.8-2018 standard which requires the swirl angle to be less than 5°, and therefore optimization of the floor splitter plate design is needed. The simulation of flow in a full-scale pump sump produced similar findings with the experimental results. From the contour and streamline plot, it was found that the immersion of the floor splitter plate has disrupted the vortical flow under the pump inlet and provided a flow straightening effect to eliminate destructive vortices and reduce swirl angle in the pump intake.
\nThe research has been funded by the Ministry of Energy, Science, Technology, Environment and Climate Change (MESTECC), Malaysia, under Science Fund grant No. SF1326 and carried out in collaboration with the Department of Irrigation and Drainage (DID), Malaysia.
\nOccupational noise is the most common health hazard that is predominant in most workplaces. In a recent survey of working adults in Canada, 42% reported being exposed to hazardous noise levels in the workplace [1]. Exposure to excessive occupational noise can cause permanent hearing loss through sensory-neural damage in the cochlea. In general, hearing is first affected in a specific range of audible frequencies (3000 to 6000 Hz) and then spreads to higher and lower frequencies. Hearing loss is often accompanied by other long-term auditory effects, such as tinnitus (ringing in the ears); increased sensitivity to loud noise; and poorer frequency selectivity (i.e., decreased ability to hear sounds in background noise) compared to individuals with normal hearing. It can also cause other, non-auditory adverse effects, the most common been the cardiovascular (e.g., changes in heart rate, increasing blood pressure). Being a stressor, noise causes also important psychological effects [2].
Noise levels in the workplace vary in level, duration and frequency content. In general, they are of high levels and are persistent for most of the work shift. They can be continuous, impulsive or interrupted. From the frequency point of view, most are of the wide band type, although they can be rich in high or low frequencies, especially if vibrations are also present in the workplace.
Reduction of the sound levels and, consequently the risk of noise induced hearing loss is the objective of every hearing conservation program in the industrial world [3].
The approach to the reduction of the risk follows several steps. The first is finding and recognizing potentially hazardous areas in the workplace. This tends to be done as a result of personal, subjective observations, the principal been difficulties in understanding speech: people ask frequently questions and answers to be repeated. Complaints of excessive noise are also important indications that the noise may be so loud as to create a health risk. This first step is usually performed through a walk-through survey. Sometimes, spot noise level measurements are also done using a sound level meter.
Once the areas with high noise levels have been found, the next step is to quantify the risk. This is done by measuring the noise exposure of individuals or groups of workers working in those areas. This procedure is known as the exposure survey.
Also, the extent of the exposed population (number of exposed persons) is also quantified to find out the magnitud of the problem.
Noise exposure is a fundamental concept in assessing the risk from high noise levels.
It is universally accepted that hearing loss occurs as a consequence of long duration exposures to high noise levels. What is usually not too clear is how long the “long duration” is and how high are the “high noise levels”. There is no, however discussion regarding that the effect is caused by a combination of both: duration and level. The concept of noise exposure combines both causes and that makes it so important. As mentioned above, in determining the risk of occupational hearing loss, measuring workers’ noise exposure is an essential part of any hearing conservation program.
It all derives from an ISO standard [4] that estimates the probability of acquiring noise induced hearing loss after being exposed to a given noise exposure level for different periods of time. As an example, after 40 years of been exposed to 85 dBA for 8 hs a day, 50% of the population will acquire an average of extra 5 dB hearing loss between 500 Hz and 6 KHz, on top of the hearing loss due to age.
On the basis of the above statement, the limit of 85 dBA has been adopted almost internationally for a workday of 8 hs.
Reference [5] lists important standards from different institutions, related to noise exposure
Noise exposure is a complex combination of sound levels a person has been exposed to and the duration of each one of those sound levels [5, 6, 7, 8]. The closer analogy is to think in terms of noise energy that enters the persons’ ears and damage the delicate organ of hearing. So, two variables are involved there: sound levels and time duration [9].
There are several concepts involved that need to be explained and defined. Their understanding is essential when dealing with this issue.
Equivalent sound level, Leq, t in dBA is the first of them. The easier way to understand it is as follows: In real life, sound levels constantly vary with time. They rise when the worker is using a power tool and diminish between operations, while changing continuously. Leq, t is a kind of an “average”, constant sound level for the entire period of exposure (working) time, encompassing all “quiet” and “noisy” periods, with the same energy of the real one. It is defined as the value of a noise of constant sound level that contains the same total A-weighted acoustical energy as the sound of interest. In other words, while the real noise is of a varying sound level, the equivalent has a constant level of the same energy.
Now is the time to clarify the meaning of the letter “t” at the end of the Leq, t. It is there to signify that the Leq in question is for the period of time the worker has been exposed to.
Here we arrive at another important point that needs to be stated: whenever Leq is mentioned, the duration of the exposure (t), should also be stated. Otherwise the Leq has no meaning. This is not too difficult to understand as per the following example: suppose we have two workers. One of them is exposed every day to 90 dBA for 4 hs. The other one is exposed also to 90 dBA, but for 8 hs. It is obvious that the effect to the hearing of the second worker will be larger. In other words even though Leq,4 of the first is equal to the Leq,8 of the second, their effects are not the same.
The numerical definition of Leq, t is as follows: ten times the logarithm (base 10) of the time integral over a stated time, t hours, of the squared A-weighted sound pressure relative to 20 μPa, divided by that time.
Noise exposure level, Lex, T, in dBA, is another important measure. This is the one used to predict noise-induced hearing loss as per [4]. It is derived from the measured Leq,t by a simple adjustment to account for the longer or shorter duration of the workday on the workers’ hearing. In other words, it answers the following question: what will be the value of Leq,t if the energy that entered the worker’s ear during t hs would enter during 8 hs. By calculating Lex,T (with capital T), Leq,t for working days of different durations can be compared directly.
The following formula converts Leq,t into Lex,T:
Where: t is the duration of the actual exposure, in hr. and
T is the normalized duration, usually = 8 hr.
As an example, if a worker is exposed to 85 dBA for four hours a day (Leq,4), his exposure for a normalized 8 hs duration will be:
If, on the contrary, he is exposed to 85 dBA for 12 hs (Leq,12), his exposure for a normalized 8 hs duration will be:
The above example shows again how two workers with the same Leq,t, have different Leq,T and, consequently, different risk of hearing loss.
Mathematically, Lex,T is defined as ten times the logarithm (base 10) of the time integral of the squared A-weighted sound pressure relative to 20 μPa for the time actually worked, divided by T hours (usually the standardized shift duration of 8 h).
Finally, it has to be stated that while Leq,t is essentially measured, Leq,T is calculated from the Leq,t value. As it will be described further, the actual measuring instrument, the dosimeter, performs both the measurement and the calculation. Both values, Leq,t and Leq,T can be read on the same device. This greatly simplifies the task of the person performing the noise exposure survey. On the other hand, it can create misunderstandings if the operator does not has clear knowledge of the difference between Leq,t and Leq,T. As mentioned above, the one that is to be used when assessing the risk of hearing loss is the noise exposure level, Leq,T.
Noise dose in % is another important measure. Although the use of the noise dose is declining lately, many instruments still allow its measurement. The concept is familiar mainly to Occupational Hygienists and commonly used when dealing with hazardous substances. The idea is quite simple: it defines the relation between the amount of a substance absorbed by a person in a given period of time (usually 8 hs) and the maximum allowed by a local jurisdiction. For example, if this limit is set to 85 dBA for an exposure of 8 hs and the actual exposure for the same period of time has been 88 dBA, then his dose will be 200%1.
The following equation allows for the calculation of Leq,t from a given dose2:
where D = dose in % for 8 h.
T = duration of the daily exposure in hours.
Lc = criterion sound level in dBA3.
For example, a dose of 100% acquired during 4 hs (using Lc = 85 dBA) will result in
Criterion level (LC) in dBA is a constant sound level which, if it continues for the criterion duration (usually 8 hs), will result in the worker’s allowable noise exposure. ISO (the International Organization for Standardization), as well as most Canadian provinces [10] and NIOSH (the USA National Institute for Occupational Safety and Health) [11] has adopted LC = 85 dBA for 8 hs.
Exchange rate is the increase (decrease) in sound level for which permissible exposure time is halved (doubled)4.ISO, most Canadian provinces and NIOSH has adopted 3 dB exchange rate. So, for instance, if a person is allowed to have Lex(8) = 85 dBA for 8 hs, he is also allowed to Lex(4) = 88 dBA for 4 hs.
There are two issues involved in the measurement of Leq,t: one is related to the instrumentation involved and the other deals with the measurement technique and procedures. Although managing the instrument itself is a relatively simple task, the measurement procedure requires basic knowledge of noise as well as practical knowledge regarding where to put the dosimeter, for how long to measure, etc. Measuring noise exposure of groups is more complex and requires some knowledge on statistics to be able to decide how many individuals to sample and for how long.
Noise exposure can be measured using regular sound level meters and integrating sound level meters. However, there is a device specifically designed to measure Leq,t. It is the noise dosimeter. In its basic version it consists of an ¼” diameter microphone connected through a long cord to a container with the battery and the electronic components of the instrument. It also includs a readout device that allows for reading of the measured Leq,t. The microphone is to be attached close to the ear of the person whose exposure will be measured. The rest of the instrument is usually worn on the belt or in the shirt pocket (see photographs in Figure 1a and b).
Dosimeters with separate microphones.
Recently, manufactures have opted for compact, small size dosimeters called Noise Badges that contain both the microphone and the microprocessor of the instrument. By having the entire instrument in a single body, they eliminate the cord that is a nuisance and also can be a workplace hazard. Measurement results can still be read on the dosimeter itself. Thay can also be transmitted via Bluetooth technology to another device with facilities for recording for future use. This is especially handy when a noise exposure survey is carried out on several workers simultaneously, while each is carrying his own dosimeter. In some models, the receiver is also a charger for the batteries of all instruments. Figure 2a and b shows Noise Badges from two manufacturers.
Dosimeters with incorporated microphones (noise badges).
There is a wide variety or instruments in the market, able to perform different measurements and calculations. They all belong to the following two basic types of dosimeters: measuring and logging.
Measuring dosimeters allow for the straight measurement of Leq,t and, eventually calculate Lex,T. Although most allow for reading the results on the instruments themselves, some others relay on a separate measurement device. This is done to keep the results visible to the operators only.
Dosimeters measure sound levels at predetermined intervals of time. Measuring dosimeters do not allow for extracting individual readings, just the final results at the end of the measurement period. Logging dosimeters, on the contrary, allow for the extraction of individual Leq,t. In such a way one can obtain the entire history of the sound levels at predetermined time intervals. The results can then be downloaded into a computing device and shown as a graph, spreadsheet, etc. By analyzing the partial data, one can follow their variation with time. Then, by knowing where the person was located at different times of the day or what kind of operation he was involved in, one can pinpoint the important noise sources or operations. Noise history is a powerful tool used for the design of noise controls in the workplace.
Another advantage of the logging dosimeters is that by studying the noise history one can determine if there have been abnormal events and then “clean” false results caused from malingering or noises not normal in the particular workplace.
Measuring Leq,t of individuals using a dosimeter is a relatively simple exercise, generally explained in the manual supplied with the instrument5. Manuals contain also information on how to care and the main precautions that have to be taken to obtain proper results.
A most important task, often overlooked, is to inform the person(s) under test the reason for testing and how it will be done. In many instances not knowing the “why” and “how” lead to malingering and falls results. Often workers suspect that the instrument will in fact transmit their conversations to the supervisor. In other instances, some individuals created artificially loud noises to show levels that do not exist in reality.
After calibrating the instrument and ensuring that the batteries have enough charge to last during the testing period, the microphone of the dosimeter is attached close to the wearer’s ear (generally on the shoulder or close by, and switched on. Then the individual is sent to perform his tasks as usual. If the task is repetitive, then the measurement is done during a couple of repetitions, only. However, when the sound levels vary during the shift or if the worker works in different places, the measurement should last for the entire shift.
As mentioned above, if the measurement has been performed for the entire shift, then Lex,T is equal to Lex,t. In other words, the daily reading is his daily noise exposure, Lex,T. If that is not the case, then the Eq. [1] (page YYY) should be used to convert the measured Leq,t in Lex,T.
In many instances, there is a need to assess a group of workers that perform identical tasks or are located in the same environment. Providing each one of them with a dosimeter is not necessary or practical. There are procedures to be followed that reduce considerably the number of instruments needed and still obtain reliable, statistically significant results6.
Noise induced occupational hearing loss is the effect on a person being exposed to high noise levels for extended periods of time. Epidemiological data, used as bases for our present knowledge of hearing loss, were derived from populations working for many years in such high noise environments [12]. This is also the origin of the equal energy theory and the 3 dB exchange rate [13].
As explained above, when the measurement period t is different from T = 8 hs, Eq. 1 is to be used,. The formula is meant for 8 hs long work day where acoustical conditions repeat day after day, month after month, for the assumed 40 active years of a person.
Presently, in many occupations, the duration of the workday is 12 hs a day with several days off to equal to 40 hs a week or 80 hs every two weeks. The question is, shall we still use Eq. 1 with T = 8 hs? No official document exists for such a situation. However, common sense indicate that since the average duration of the workday is still T = 8 hs, (the average over the 2 or the 4 weeks), Eq. 1 is still valid and shall be used.
As an example [14], the total of hs worked by the musicians at the National Ballet of Canada is 350 hs. Therefore, the average Leq,t during their rehearsals/performances was corrected using Eq. 1 as follows:
Where t = 350 are the actual annual number of hours worked and.
T = 2000 the number of work hours in a year.
We do not really know what happens to ears exposed to 12 hs a day, for a 40 hs week. Nor we know about yearly exposures of less than 2000 hs, that is the average exposure resulting of 8 hs a day, 40 hs a week. We can only assume that the equal energy principle can be extended to cover exposures of different durations.
Using the equal energy principle, one can calculate exposures of different workday duration too. For example, if a worker whose workday is 8 hs and whose exposure measured for 5 hs was Leq,5 = 85 will be.
However, if his workday is t = 12 hs, then
In the case of temporary worker, that performs 350 hs a year, it will be t = 350 hs, T = 2000 hs and Eq. 1 will be
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