Patrizia

BUILDING ACOUSTICS

Vibration transmission of asymmetric T form junctionsF. Reis, Laboratory of Building Acoustics,

Budapest University of Technology and Economics, Budapest, Hungary

The paper explains the asymmetric T form junction, describes its occurrence and determines a method by which the acousticalcharacteristics can be determined, fitted to existing data of symmetrical T form junctions.

EXPLAINING THE PROBLEMAnalysing the plans of existing building several

forms of structural junctions can be collected. Some ofthem are solved, others have been out of the scope ofthe worked-out predicting procedures. One of theunsolved junctions is the asymmetric T.

T form junctions mostly occur at facades: thecrossing element of the T is the facade wall, theperpendicular stem corresponds to the inner partitionwall. The flat facade results the straight crossingelement of T. In real building where the facade wall iscontinuos the T form junction is symmetric. Howevermodern architecture is rich in using complex formalsolutions on the facades and in planning ground planesthe form of which is not simple rectangular. Even thefacade of an L shape building is not plain. AsymmetricT form junction can be found at junctions, where thefacade is not plain and a partition wall is coupled tothe facade wall. This is illustrated on Fig. 1.

The analyse and the results, introduced here arerelated to the simplified prediction of field weightedsound reduction index: only weighted and averagedterms and quantities are considered.

ACOUSTICAL CHARACTERISTICS OFASYMMETRIC T JUNCTIONFrom the point of predicting filed sound

reduction index the acoustical characteristics ofjunctions are the surface average vibration velocitylevel difference, Dvij or the vibration reduction index,Kij . The relation of the structure-borne powerreduction index Rij to Dvij is given in eq. (1) [1], whereindex i corresponds to the excitation side, index jcorresponds to the receiving side of the junction, f( )expresses that the relation between Rij and Dvij dependsalso on the data of the materials and sizes.

( )( )sizestamaterialdafDR vijij ;lg*10+= (1)

An important characteristics of Kij is summarised ineq. (2), based on [2], originally established also in [3]:

( ))(lg tamaterialdafRK ijij += (2)

Both equations lead to the same conclusion: thedifference between Rij and Dvij or Kij depends ongeometrical and material data, and doesn’t dependon the form of the junction.

The steps to determine the structuralcharacteristics of asymmetric T junction are thefollowing:1. Dvij or Kij of the symmetric T form junction are

given, known for eg. in [4] or [2].2. The differences according to (3) and (4) have to

be calculated for the reliable combinations ofmaterials and thickness:

vijijij DRdRD −= (3)

ijijij KRdRK −= (4)

4. Regression curves have to be fitted to thecalculated differences, a linear function of therelation of the specific masses for Dvij, and asecond order power function for Kij,.

5. The bending wave structure-borne reductionindex of asymmetric T form junction ρAij has tobe calculated, based on [5] for the reliablecombination of materials, and thickness. Fig 2.shows the symmetric ( left side ) and asymmetric( right side ) T form junction schematically, withthe numeration of the elements.

6. The vibration level difference and the vibrationreduction index of the asymmetric T formjunction, Davij and Kaij can be calculatedaccording to (5) and (6) for the discrete values ofthe relation of the specific masses:

ijaijij dRDRDa −= (5)

ijaijij dRKRKa −= (6)

7. The discrete numerical values of Davij and Kaij

have to be approximated by regression functions,the same type as the original ones.Fig 3-8 presents a set of examples for the results

of the procedure described above. In the examplesi=2, j=1 and 4.

CONCLUSIONSAs the examples show the procedure described

above in steps 1-7 is suitable to determine thestructural characteristics of asymmetric T formjunctions.

The Dvij values depend on the direction ofpropagation, this can be observed too in the graphs.Despite of the different data K21 and K24 are almost thesame: Kij is independent of the direction. In the case ofasymmetric T form junctions the derived values for

Dvij and for Kij differ from the published ones forsymmetrical T. This is the reason to introduceasymmetric junctions to the set of the existing ones.

REFERENCES[1] Khilman, T.: Transmission of structure-bornesound transmission in buildings. National SwedishInstitute for Building Research, 1967[2] EN 12345-1:2000: Building Acoustics - Estimate

the acoustical performance of buildings from theperformance of the elements – Part 1: Airborne soundinsulation between rooms[3] Gerretsen, E.:Sound transmission betweendwellings. Applied Acoustics . Vol 12, No 6, 1979 [4] W. Fasold, et al: Taschenbuch Akustik VEBVerlag Technik Berlin 1984. Fasolf, W.: 6. Bauakustik[5] Heckl, M, Cremer L.: Structure Borne SoundSpringer Verlag Berlin 1988.

Fig 1.Examples of existing building

ROOM ROOM

with asymmetric T junctions

ROOM ROOM

Fig 2.Symmetric and asymmetric T junction

1 4

2

1 4

2

Fig 3. Symmetrical T 2->1 Dv(2-1)

- 1 0

-5

0

5

1 0

1 5

2 0

2 5

3 0

0 .1 1 1 0

m " c r / m " l o

Dv(2

-1)

R(2

-1)

dB

R (2-1) calc

Dv(2 -1 ) g iven

dRD(2 -1 )=R(2 -1 ) -Dv (2 -1 )

d R D ( 2 - 1 ) r e g r

Fig 6. Symmetrical T 2->1 K(2-1)

-20

-10

0

10

20

30

40

0.1 1 10

m"cr /m" lo

K(2

-1)

R(2

-1)

dB

R ( 2 - 1 ) c a l c

K ( 2 - 1 ) g i v e n

d R K ( 2 - 1 ) = K ( 2 - 1 ) - R ( 2 - 1 )

d R K ( 2 - 1 ) r e g r

Fig 4. Asymmetrical T 2->1 Dv(2-1)

-10

-5

0

5

10

15

20

25

30

0.1 1 10

m"cr /m" lo

Dv2-1

R(2

-1)

dB

R ( 2 - 1 ) a c a l c

d R D ( 2 - 1 )

D v ( 2 - 1 ) a

D v ( 2 - 1 ) a r e g r

Fig 7. Asymmetrical T 2->1 K(2-1)

-20

-10

0

10

20

30

40

0.1 1 10

m"cr /m" lo

K(2

-1)

R(2

-1)

dB

R ( 2 - 1 ) a c a l c

d R K ( 2 - 1 )

K ( 2 - 1 ) a

K(2 -1 )a regr

Fig 5. Asymmetrical T 2->4 Dv(2-4)

-10

-5

0

5

10

15

20

25

30

0.1 1 10

m"cr /m" lo

Dv(2

-4)

R(2

-4)

dB

R ( 2 - 4 ) a c a l c

d R D ( 2 - 1 ) = d R D ( 2 - 4 )

D v ( 2 - 4 ) a

D v ( 2 - 4 ) a r e g r

Fig 8. Asymmetrical T 2->4 K(2-4)

-20

-10

0

10

20

30

40

0.1 1 10

m"cr /m" lo

K(2

-4)

R(2

-4)

dB

R ( 2 - 4 ) a c a l c

d R K ( 2 - 4 )

K ( 2 - 4 ) a

K(2 -4 )a regr

Structure-borne sound of equipment in buildings -characterisation and application of the source strength

for washing machines as example

E. Gerretsen

TNO TPD, P.O. Box 155, NL-2600 AD Delft, The Netherlands

The prediction models for building acoustics in EN 12354 so far relate the airborne and structure-borne sound transmissionbetween rooms. An important acoustic aspect remaining to be treated in these series of models is the airborne and structure-borne sound as caused by machines and equipment in buildings. Suggestions for possible approaches have been presented inearlier papers. The largest problem with this topic is the adequate and yet practical characterisation of machines and equipmentas a source of structure-borne sound. Two of the suggested possibilities to measure and express the structure-borne sourcestrength will be discussed and illustrated with results from measurements with washing machines as an example of equipmentin buildings.

INTRODUCTION

For any type of prediction model for the equipmentsound levels in buildings due to structure-borne soundexcitation of the structure by equipment, installationsor installation parts, it is essential to have input data onthe structure-borne source strength of the equipment. Itis thus important to consider feasible approaches on anengineering level to express and measure such data.A feasible approach is to consider the equipment as asingle force source (equivalent force level) [1], another- more general - approach is to use the characteristicstructure-borne sound power [2]. These approacheswill be compared for washing machines as structure-borne sound source. Also indications will be given ofthe appropriate quantity to express the soundtransmission through the building in those cases, inline with the existing models for airborne and impactsound transmission in EN 12354-1&2 [3].

EQUIVALENT FORCE LEVEL

For various sources the assumption of a low sourceimpedance compared with the usual floor impedancein buildings seems rather realistic. In that case theforce level describes sufficiently the strength of thesource at a contact point. However, normally a sourcewill have several contact points and it can bequestionable if the complete source can than berepresented by one equivalent force level. To studythis aspect measurements were performed withwashing machines (centrifuging operation mode) on aconcrete floor (100 mm concrete) in a test facility.

The equivalent force of the washing machine isdeduced by a substitution method, assuming variousfase relations ϕ between the four feet. The transfermobility Yij is measured from the position i of the fourfeet (M=4) to the position j of four accelerometers(N=4). Then with the running machine the vibrationlevels are measured at the four positions j. From thisthe equivalent force Feq follows assuming random fase,in-fase or anti-fase between pairs of feet.In case of random fase this can be simplified to theratio of the average squared velocity to the squaredabsolute value of an average transfer impedance. Sucha quantity would be fit for a general characterisation ofthe measurement facility, facilitating measurements.

In figure 1 the results are given for the average of 8machines with the different fase assumptions; the A-weighted levels for random fase vary from 120 tot 128dB(A). The largest difference occur at low frequencieswith the assumption of anti-fase between two pairs offeet (rocking machine). The only measure to decidewhich approach is most realistic is the resulting soundpressure level in the room below the test floor. Itshows that the variations in the difference betweenforce and pressure is largest with anti-fase (σ∆=7dB(A)) and about equal in the other cases (σ∆ =3 to 4dB(A)). Since this difference in octave bands is thesmoothest for the random assumptions that seems tothe most realistic and practical assumption.

)1(1

11

2

1

22

�

�=

=

=N

j M

i

jij

jeq

ieYM

vN

Fϕ

FIGURE 1. Equivalent force level, average of 8washing machines with different fase assumptionsbetween feet.

CHARACTERISTIC POWER LEVEL

A more general method to present the source strengthis the characteristic structure-borne sound power Wsc.which includes also the influence of the sourcemobility. In its simplest form it can be deduced fromthe (equivalent) force by

The actual injected power in a given situation isrelated the source mobility and the mobility of themounting structure, the coupling term. In a formerstudy the source mobility for some washing machineshave been determined. The real part is presented infigure 2 and rather constant for these machines. Theimaginary part shows a mass-spring behaviour.Assuming these data to be relevant for this type ofmachines in general, the characteristic structure-bornesound power level LWsc in dB re 1 pW is thus 31 dBlower than the force level in figure 1. For known typesof sources the characteristic power can be deduced inthis way from force levels and typical source mobility.For other sources the determination requires also theactual measurements of the mobility.

FIGURE 2. Real part of the source mobility ofwashing machines; different feet (thin lines) and trend(thick line).

STRUCTURE-BORNE SOUNDTRANSMISSION

The description of the sound transmission should bebased on the injected power. For a force source thatfollows from the real part of the floor mobility underthe assumption of a low source impedance. In the moregeneral approach that is given by the coupling term,for the one dimensional case as in (2) given by:

With the source mobility as measured it shows that onconcrete floors the injected power only depends on thefloor mobility - and the force source description wouldthus be adequate here. But on a wooden floor themobility of both floor and source become important.

REFERENCES

1. Gerretsen, E., Prediction model for sound transmissionfrom machinery in buildings: feasible approaches andproblems to be solved, Proc. Inter-noise 2000, Nice;2000, pp. 3345-3350.

2. Moorhouse, A.T. & B.M. Gibbs, Relationship between thecharacteristic power of structure-borne sound sourcesand their emission when installed, Proc. Euronoise '98,Munich; 1998.

3. EN 12354, Building Acoustics - Estimation of acousticperformance of buildings from the performance ofelements, part 1&2, 2000.

100

110

120

130

140

150

160

170

180

16 31,5 63 125 250 500 1000

fre que ncy [Hz]

L Feq

[dB

re 1

N

]

random

in fase

antifase

)2()Re(2sourceeqsc YFW =

-50

-40

-30

-20

50 100 200 400frequency [Hz]

10 lg

ReY

[dB

re 1

s/k

g]

)3()Re(/)Re(

2

sourceflr

sourceflrscinj

YY

ZYWW

+=

Determination and Use of Transfer Functions to describeStructure-borne Sound Transmission Caused by

Equipment and Installations

H-M. Fischer

Fachhochschule Stuttgart/University of Applied Sciences, Department of Building Physics, Schellingstrasse 24,D-70174 Stuttgart, Germany

For the characterisation of structure-borne sound transmission in buildings so called global transfer functions can be used todescribe the complete transmission path between a certain excitation point and a receiving room. The paper explains how toreceive those transfer functions experimentally by means of a reciprocal measuring method and how to use them for a simpleprediction of the expected equipment noise. Additionally it can be shown that the used reciprocity method advantageously canbe used to describe structural sound sources which can be characterized as point sources. Examples gained in differentbuildings will demonstrate the use and interpretation of the transfer functions.

INTRODUCTION

Problems due to technical equipments in buildingsoften are caused by generation and transmission ofstructure borne sound. One of the most prominentquestions in building acoustics is how to describestructural sound sources and how to calculate soundtransmission with respect to a prognosis of the soundpressure levels to be expected in a certain buildingsituation. The following contribution will concentrateon applications, where the structural sound sources canbe regarded as point sources (or small sources) and thebuilding constructions as heavy constructions. If theinput impedance of the excited structure is greatcompared with the source impedance the structuralsource may be described as force source. These limitations restrict the application of thediscussed prediction method in general butnevertheless it covers a wide range of present technicalinstallations in many buildings. In this contribution the experimental determinationof source and transmission characteristics will beattributed to measurements based on the reciprocityprinciple. This principle is well known in acoustics butnot really introduced to building acoustics. In [1]Feldmann and Buhlert have shown the basicapplication of reciprocity to gain source andtransmission data for the purpose of building acoustics.Neverthelesss the described methods did not findentrance to practical applications. It is the aim of thispaper to demonstrate the possibilities of this helpfullinstrument for characterisation and prediction ofinstallation noise.

DESCRIPTION OF TRANSMISSIONPATHS

For predictions of the radiated sound pressure level thetransmission path could be calculated step by step ordescribed in total by a global transfer function whichincludes the complete transmission path from theexcitation point on a certain structure to the radiationin a given receiving room. The first approach allowspredictions more in detail but requires knowledge ofall input data. The second approach, which will beoutlined here, allows predictions only for thosesituations, where an appropriate transfer function isavailable, but it enables a simple prognosis. In generalthe global transfer function can be described as the(frequency dependend) ratio of the exciting point forceF to the total radiated sound power Wtot. In [1] thisratio is given in the following relation and called thestructural sensitivity �F.

2tot

2F FW

kρcα �� (1)

(k: wavenumber, �c: characteristic impedance).�F can be measured directly for any transmissionsituation, if the force spectrum is known. Analternative method to determine �F indirectly is givenwith respect to reciprocal relations. In this case theexperimental determination of �F is based on theexcitation of a reverberant sound field (with soundpressure p) in the receiving room and the measurementof the velocity v of the vibrating structure at the sameplace as for force excitation in the previous case. It canbe shown [1] that the transfer function now can be

expressed by

2

222

F pv

4cρα �

�

� (2)

This reciprocal relation advantageously can be usedfor the experimental determination of the neededtransfer-functions: no special structure borne soundsource for excitation is required, the measurement offorces is not necessary and the velocities can bemeasured directly at the mounting points of thesources of interest.

DESCRIPTION OF STRUCTURALSOUND SOURCES

The reciprocal method also enables an indirect methodto determine unknown forces of structural sources.From eq. (1) follows

F

tot2

2 WkρcF

�

�� (3).

In this case the radiated sound power of the structuralsource has to be measured in a receiving room and �Fmust be known for the chosen transmission situation.This may be a laboratory or any other buildingsituation. Thus the forces can be determined under in-situ-conditions of the source.

PREDICTION OF SOUND PRESSURELEVELS

From eq. (1) directly follows the procedure how topredict the radiated sound power in a giventransmission situation (caracterized by �F) and for agiven structural source (characterized by its forcespectrum):

2Ftot F

c²kW ���

�� (4).

For practical applications this prediction procedurerequires the sufficient availability of transfer functions�F for the most important transmission situations andof force spectra of the sources. These data have to becollected and could be presented in a suitablecatalogue. Some examples of transfer functions areshown in the following chapter. The predictionprocedure is valid if the source characteristics areindependend from the excited structure. This holdstrue if the sources can be considered as force sources .Thus the proposed prediction method mainly isapplicable for smaller technical equipment (e.g.sanitary installation) in combination with solidconstructions.

EXPERIMENTAL RESULTS

Meantime a number of transfer functions and sourcespectra have been collected by means of the describedreciprocity measurements. They will serve as the basisof a data catalogue for simple predictions. Someexamples are given in figures 1 and 2.

REFERENCES

1. Buhlert, K, Feldmann, J.: Ein Meßverfahren zur Bestim-mung von Körperschallanregung und -Übertragung,Acustica 42 (1979), 3

1,0E-09

1,0E-08

1,0E-07

1,0E-06

1,0E-05

1,0E-04

1,0E-03

50 100 200 400 800 1600 3150

frequency [Hz]

Alp

ha F

horizontal diagonal

1 ,0 E -0 9

1 ,0 E -0 8

1 ,0 E -0 7

1 ,0 E -0 6

1 ,0 E -0 5

1 ,0 E -0 4

1 ,0 E -0 3

5 0 1 0 0 2 0 0 4 0 0 8 0 0 1 6 0 0 3 1 5 0

fre que ncy [H z ]

Alp

ha F

h o llo w b rick s , m ''= 1 80 k g /m ²C a -S i b locks , m ''= 24 3 kg /m ²au toc l. a e ra ted conc re te , m ''= 27 3 kg /m ²con c re te b lo cks , m '' = 385 kg /m ²

FIGURE 2. Structural sensitivity �F of masonrywalls of different materials for horizontal transmission

FIGURE 1. Structural sensitivity �F of a masonry wall(Ca-Si-blocks, m' = 243 kg/m2) for horizontal anddiagonal transmission in a solid building

The Effect of Wall Ties in External Cavity Walls on theAirborne Sound Insulation of Solid Separating Walls

R. Hall, C. Hopkins and P. Turner

Acoustics Centre, BRE, Garston, Watford, WD25 9XX, UK

For thermal insulation, external walls in the UK are typically cavity walls with a brick outer leaf and a masonry inner leaf.Flanking laboratory measurements show that the airborne sound insulation of a solid masonry separating wall is affectedby the wall ties used in the external masonry cavity flanking walls in addition to the masonry inner leaf material.Therefore, there can be conflicts between thermal insulation and sound insulation. This paper presents measurement datathat demonstrate the effect of the external cavity wall construction on sound insulation with solid masonry separatingwalls, in particular, the effect of dynamically stiff wall ties on the sound insulation of a solid masonry separating wall withlightweight flanking walls.

INTRODUCTION

Exterior cavity walls with large cavities allowthicker thermal insulation in the cavity and thetypical UK cavity width will increase from 50mmto 100mm. For structural reasons, stiffer wall tiesmust be used with 100mm cavities than with 50mmcavities. The data presented in this paper wereobtained from an investigation to discover whetherusing stiffer wall ties leads to an increase inflanking sound transmission and lower airbornesound insulation of separating walls. The BREflanking laboratory was used for the measurements.The test construction is illustrated in Figure 1 anddescribed in Table 1. A mineral fibre cavity stopwas placed in the cavity at the separating wall.Vertical twist ties having a measured dynamicstiffness1,2 (k100mm) of 43.4MN/m were used toconnect the inner and outer leaves of the exterior

masonry cavity flanking wall. The effect of the wallties was assessed by measuring the airborne soundinsulation and vibration reduction index, (a) withthe wall ties fixed in both leaves of the exterior walland (b) with the wall ties disconnected from theouter leaf.

Wall 413

Wall 2

Wall 1

Wall 3

Room 2 Room 1

FIGURE 1: Plan view and numbering system forthe test constructions.

Table 1: Test construction details

Element Material Thickness(mm)

Surfacedensity(kg/m2)

Measuredlongitudinalwavespeed

(m/s)

x (m) y (m)

Separating wall 1

Dense aggregateblocks with 13mm

plaster (both sides)

215 440 2880 4.90 2.34

Inner flanking walls 2 & 3

Lightweightaggregate blocks with

13mm plaster (one side)

100 150 2900

4.15 (Wall 2)

4.25 (Wall 3)

2.34

Outer flanking wall 4 Brick 100 170 2500 8.58 2.77

MEASUREMENT RESULTS

The ISO 140-4 airborne sound insulation data forthe separating wall with and without wall ties areshown in Figure 2. Disconnecting the wall tiesincreased the DnT values by 2.3dB – 6.5dB in thefrequency range 50Hz – 315Hz causing a 2dBincrease in DnT,w.

25

30

35

40

45

50

55

60

65

70

75

50 100 200 400 800 1600 3150

Third octave frequency band (Hz)

Dire

ctio

n av

erag

e D

nT

(dB

) W ith wall ties 55(-2;-6;-1;-8)

W ithout wall ties 57(-2;-6;-1;-7)

FIGURE 2: Airborne sound insulation with andwithout wall ties. DnT,w(C;Ctr;C50-5000;Ctr,50-5000)

Vibration reduction index measurements were takenaccording to the draft standard prEN ISO/DIS10848-1. Measured K23 data are shown in Figure 3,where K23 corresponds to vibration transmissionbetween inner flanking walls 2 and 3.

The most important part of the frequency range isthat from 100Hz to 630Hz because the adversedeviations in the airborne sound insulation ratingfor the single-number quantity occur in thesefrequency bands.

0

2

4

6

8

10

12

14

16

18

20

100 200 400 800 1600 3150

Third octave frequency band (Hz)

Vibr

atio

n R

educ

tion

Inde

x K

ij (

dB)

K23 (W ith wallties)

K23 (W ithoutwall ties)

FIGURE 2: Vibration reduction index K23 with andwithout wall ties.

CONCLUSIONS

The measured data demonstrate the significance offlanking transmission at low frequencies via wallties with high dynamic stiffness in external masonrycavity walls. In the 100Hz to 630Hz frequencyrange, flanking transmission via the wall ties usedin this test construction reduced DnT,w by 2dBcompared with the situation where the ties weredisconnected from the outer leaf of the cavity wall.

The estimation model for airborne sound insulationdescribed in BS EN 12354-1 is generally suited tosolid separating walls, but does not explicitlyaccount for flanking transmission via wall ties inacross the external wall. The results from thisproject demonstrate the potential importance of thisflanking path.

ACKNOWLEDGMENT

This research was funded by the UK Department ofEmployment Transport and the Regions (DETR).

REFERENCES

1. C. Hopkins, R. Wilson and R.J.M. Craik. AppliedAcoustics 58, 51-68 (1999).

2 R. Hall and C. Hopkins, Dynamic stiffness of wallties used in masonry cavity walls: measurementprocedure, IP3/01, ISBN 1 86081 461 1, CRC Ltd.PO Box 202, Watford, WD25 9ZW, UK, (2001).

Systematic Comparison of Sound Insulation Measured in situwith Values Calculated According to EN 12354 Yields

Confident Data of the Acoustical Properties of BuildingElements

C. Simmons

J&W Akustikbyrån, SE-415 26 Gothenburg, Sweden

In a project on a new Nordic building element database, input data to EN 12354 are established from laboratory data, in situdata and theoretical models. Each case is normalised, i.e. calculated according to EN 12354 to correct for the flankingconditions and structural reverberation. The input data of the elements are then fitted to measurements such that the meandeviation between measured and calculated values are minimized. The standard deviation is used to estimate the confindenceof predicted values. Two applied cases are presented. One where the performance of a 180 mm concrete slab is calculated forvarious flanking conditions to find building plans where the requirement R'w+C50-3150 � 56 dB may be fulfilled. Secondly,input data for a floating floor is established and the performance of the assembly of hollow core conrete slabs and the floatingfloor is compared with in situ measurements. A 3 dB prediction margin to a required performance appears reasonable.

INTRODUCTIONThe standards EN 12354 (1) were developed for

the purpose of estimating sound insulation in buildingsfrom acoustic input data for building elements. J&Wand Delta Acoustics & Vibration cooperate on aNordic database project where input data to EN 12354are established from data taken in the laboratory or insitu or theoretical models. There are advantages anddrawbacks with all the approaches mentioned.

- Laboratory measurements may be “exact” in alimited sense (if structural losses at the testinglaboratory are determined), but the data obtained areoften regarded by consultants as being “impractical”,“unrealistic” etc because they do not take into accountlow frequency room modes, flanking transmission,structural absorption and building practice in situ.

- Theoretically derived input data may be biasedwith respect to errors in the model or the material data,but random errors are often small.

- Data taken in situ to find suitable input data forbuilding elements from field measurements requiressome processing. Otherwise a prohibitive data scatteron "similar" partitions will result as is shown in figure1. They confirm our negative experience from earlierdatabase projects. In a Nordtest project (2), a methodwas derived and tested thoroughly that reducesuncertainty in EN 12354 predictions. Measurementdata in situ from many types of building elementstypical for the Nordic countries were collected. Eachfield case was normalised, i.e. a calculation was madeof the building according to EN 12354 (using the

CADBA® software). The input data of the buildingelements were fitted as to minimize deviationsbetween measured and calculated values. This proce-dure was repeated for each of the measurements. Thestandard deviation of all cases was used as an indi-cation of the confidence in predictions. The 90%confidence interval in predicting R'w of heavy elem-ents was then about 4 dB. In the J&W / Delta databaseproject, only data from well documented buildingobjects and satisfactory measurement condititionswere included to reduce the uncertainty. For heavyelements, such as hollow core concrete slabs, thecalculation model derived by Delta adapts to Annex Bof EN 12354 with some modifications. The reductionof airborne sound insulation at the coincidencefrequency appears to be higher than expected fromtheory. Impact sound of thin hollow core concreteslabs increases at higher frequencies due to reducedstructural losses. These models of concrete slabs areused in the calculations described below.

EXAMPLES OF THE INFLUENCE OFFLANKING TRANSMISSION AND

STRUCTURAL LOSSESSix cases of vertical transmission through a 180

mm massive concrete slab are calculated and presentedin figure 1 below. The element input data for the slab(as would be measured in the laboratory) are indicated.The flanking transmission is calculated according toEN 12354 for some typical floor plans of dwelling

houses. Three parameters are varied: 1) two sizes ofroom (bedroom 10.5 m2 (BR) and living room 36 m2(LR) are used which change the relation between areaand perimeter of the common slab. 2) the outer wallsare taken as either heavy concrete elements (HF) orlight-weight plasterboard walls (LWW). 3) the roomsare placed either in a corner (CO) or in the middle of abuilding (CE). The scatter in results is in the order of 8dB which means that it is not sensible to use fieldresults to predict the performance of the buildingelement alone unless they are “normalized” withrespect to structural losses.

Normalized impact sound level (L n), dB

15

20

25

30

35

40

45

50

55

60

65

50 80 125

200

315

500

800

1250

2000

3150

5000

Frequency, Hz

COLRHF

COLRLWF

CELRLWW

COBRHF

COBRLWF

CEBRLWW

L 'nw+Ci, 50-2500

FIGURE 1. Impact sound insulation of a 180 mm concreteslab, as measured in the laboratory (bold line) and with sixcases of flanking transmission. Legend: CO corner room, CEcentral location, LR living room, BR bedroom, HF heavyfacades, LWW light weight facades.

FLOATING FLOOR PERFORMANCESome examples are shown below, where the mean

performance of a typical floating floor is estimatedfrom measurements in 7 buildings. This wooden floor(22 mm, 15 kg/m2) is mounted on resilient pads onhollow core concrete slabs (300 kg/m2). Theassembled thicknesses of the floors are 70 or 150 mm.Figure 2 show the mean impact sound improvementand the standard deviation of performance. Using theseinput data, R'w+C50-3150 and L'nw+Ci,50-2500 were

estimated correctly in 7 buildings within [0;-3] dB ascompared to measurements in situ.

Impact sound reduction: mean and standard dev. (dB)

0

10

20

30

40

50

60

50 63 80 100

125

160

200

250

315

400

500

630

800

1 k

1,25

k

1,6

k

2 k

2,5

k

3,15

k

Frequency (Hz)

70mm, DL

70mm, DL (std)

150 mm, DL

150 mm, DL (std)

FIGURE 2. Impact sound reduction of two types of floatingfloor (70 mm and 150 mm height) estimated from 7 in situmeasurements. (�) bias error (- -) standard deviation.

The scatter at high frequencies for the 70 mm heightare probably related to less confident building practice.The 150 mm floors are less sensible to structure bornesound bridges from ductwork in the air gap.

ACKNOWLEDGEMENTSI thank Dan B. Pedersen at Delta Acoustics and

Vibration and Heinrich Metzen at Isover AG.

REFERENCES1. EN 12354. Building Acoustics - Estimation of acoustic

performance of buildings from the performance of elements -Parts 1 and 2.

2. D.B. Pedersen. Nordtest technical report 425 1998.

3. BASTIAN ® version 1.1. G+H Isover 2000.

On Using Multiple Kij’s in the EN12354 Acoustics PredictionModel to Represent Excess Attenuation in Flanking Surfaces

T. RT Nightingalea and I. Bosmansba Institute for Research in Construction, National Research Council, Ottawa, Ontario, Canada K1A OR6

bCSTB, 24, rue Joseph Fourier, 38400 Saint-Martin-d'Hères, FranceEN12354 is a standardised method to compute the apparent sound insulation of building assemblies formed by wall and floorelements that are assumed to be heavy, monolithic, homogeneous and moderately damped. However, most large walls and floorsare not perfectly homogeneous, they are made by joining together a series of smaller more manageable pieces, (e.g., precastconcrete, plasterboard, glass panes, etc.). Depending on the mechanical properties of the joints between panels, there may besignificant excess attenuation, which in most cases indicates the surface must be treated as a series of connected panels. Thissummary paper shows that EN12354 is based on simplified SEA and can be extended to account for excess attenuationintroduced by junctions between the elements. However, the effective Kij necessary to account for the excess attenuation is notsimply the sum of the Kij’s for all junctions, as this leads to significant errors.

INTRODUCTIONThe EN12354 [1] prediction models allow for

one junction or attenuating element in the flanking pathwhich permits estimation of the flanking soundreduction for most paths. The path 1453 of Figure 1Ais one example. Recently, situations have beenpresented [2], similar to that shown in Figure 1B forpath 14653, where a small element (shown as 6) couplesa larger element (5) to the junction. In this case thereare two junctions (4-6 and 6-5) in the path each causingvibration attenuation. It has been suggested [2] thatexcess attenuation due to multiple junctions could berepresented by an effective vibration reduction index,K*ij, which is simply the sum the Kij’s for the two, ormore, junctions of the flanking path.

1

2

3

4

R1453

5

R14653

53

14

62

A B

FIGURE 1. Two rooms 1 and 3 separated by wall 2. InFigure 1A the flanking walls 4 and 5 do not have a couplingelement, while in Figure 1B they are coupled by element 6.

This paper provides a derivation based on theprinciples of SEA to show that for the case of multiplejunctions it not possible to sum the vibration attenuation

of all the junctions in the flanking path as, in general,this will lead to serious errors.

THEORYUsing SEA [3] the sound reduction index

(SRI) for an arbitrary transmission path is,

���

����

����

�

����

��

EA

E

E

D

D

C

C

B

B

AABCD AV

SVEE

EE

EE

EER log10log10 (1)

where E is the energy in the subsystem, V is the roomvolume, S is the partition area separating the two rooms,A is the absorption in the receive room, and thesubscript identifies the element in the transmission path.

Using equation 1, and a number ofapproximations and assumptions, the EN12354expression for the SRI of a flanking path involving asingle junction (equation 25 in Ref. 1) can be obtained.A detailed discussion is beyond the scope of thissummary paper and will be presented elsewhere; insteadthe procedure is summarised as follows. The energystored by each subsystem is expressed in terms of theratio of the total loss factor to coupling loss factor, andusing reciprocity with several substitutions, theexpressions for path 1453 can be written as,

���

����

���

��

����

��

��

4554

54

54

2254

1453 log10log52 ��

��

SSSRR

R (2)

while the corresponding expression for path 14653involving two junctions is,

���

����

���

��

����

��

��

56654664

6564

54

2254

14653 log10log52 ����

����

SSSRR

R

(3)where Ri is the resonant sound reduction index for theelement i when measured in-situ. By inspection ofequations 2 and 3 it can be seen that the basicformulations are identical excepting the last terms whichaccount for junction attenuation. Thus, it would appearthat multiple junctions can be accommodated bymodifying the junction term. It must be noted that terms

involving element 6 appear only in the description ofthe junction, consequently element 6 is assumed not toradiate energy. Thus, element 6 is referred to as being acoupling element.

The junction terms of equations 2 and 3 can bedescribed by the velocity level difference, VLD, of theconnected elements and their mass, using,

��

�

�

��

�

���

��

�

�

��

�

�

j

iijv

ij

j

MMD log10log10 ,

�

� (4)

Reciprocity that was used in writing equation 2 greatlysimplifies the junction terms as they reduce to thedirection averaged VLD, ijvD , , between the connectedplates i and j. The term for a single junction becomes,

45,54,45,

4554

54

2log10 v

vv DDD

��

����

����

�

��

�� (5)

while the expression for two junctions becomes,

65,46,

65,56,64,46,

56654664

6564

22log10

vv

vvvv

DD

DDDD

��

��

����

�

����

�

����

����

(6)

EN12354, expresses the vibration attenuationat the junction in terms of a vibration reduction index,Kij, which is the direction averaged VLD normalised tothe length of the junction, lij, and an “equivalentabsorption length”, a, of each connected element,

ji

ijoijvij aa

llDK log10, �� (7)

For framed elements, ones with a high internal lossfactor, or ones which are considerably lighter than theelements to which it is connected, EN12354 assumesthat a, can be approximated by the surface area of theelement, S, divided by unit length, lo. Using this andsubstituting equation 7 into 5 gives the EN12354 (Part1, equation 25) result for a single junction,

���

����

���

��

45

245

541453 log10

2 llSKRRRo

(8)

while the corresponding expression for two junctions is,

���

����

���

���

����

���

��

65

665

46

246

5414653

log10

log102

llSK

llSKRRR

o

o (9)

For the sake of simplicity in the followingdiscussion we shall assume that the length of all thejunctions are the same, l, which is often the case forwalls. If the expression applicable to a single junctionas given in EN12354 (i.e., equation 8 above) were to beextended to allow for two junctions with a coupling

element then the effective junction vibration reductionindex, *

45K , would be,

���

����

����

ollSKKK 6

6546*45 log10 (10)

DISCUSSIONEquation 10 indicates that in general the Kij’s

can not simply be summed since there is an additionalterm involving the area of coupling element 6 and thejunction length. In most cases, element 6 will be arectangular with the length of one side equal to thecoupling length l, and the other side equal to w.Consequently, the error can be approximated by:

� �wll

S

o

log10log10 6 ����

����

� (11)

The assumption of summing Kij’s may underestimatethe SRI when the width, w, is greater than 1 metre andoverestimate when the width is less than 1 metre.

Although it might be tempting to sum the Kij’sand apply a correction term equal to equation 11, itshould be stressed that the expressions given inequations 9, 10 and 11 are only valid when element 6satisfies the conditions necessary for a SEA platesubsystem. This requires that element 6 support enoughmodes in both principal directions so that the modaloverlap factor exceeds unity, and that the vibration bedominated by the reverberant field (that is the responseshould be reasonably diffuse).

When element 6 is a narrow strip, thesefundamental and necessary conditions are most likelynot satisfied and any form of an effective Kij should notbe applied at all. In this situation, one should treat bothjunctions and the intermediate element as a singlecoupling element characterised by a single Kij.

In the derivation of equations 9, 10 and 11, itwas assumed that the sound radiated by element 6would be small with respect to element 5 which iscoupled to it. However, for situations where element 6satisfies the conditions of an SEA subsystem, it is verylikely that radiation will be significant and it may benecessary to compute the SRI of paths 1463, in additionto 14653.

The derivation shows that EN12354 is basedon simplified SEA where there is only one junction,consequently care must be given to ensure that allelements satisfy the requirements of an SEA subsystem.

REFERENCES[1] Anon, EN12354, Estimation of acoustic performance of buildingsfrom the performance of elements, Parts 1 and 2, 1997.[2] Schumacher, Rolf, and Sas, Bernd, Journal of BuildingAcoustics, 6, No.3/4, pp. 309-340 (1999).[3] Craik, Robert J M, Sound transmission through buildings usingstatistical energy analysis, Gower Publishing Ltd., 1996, pp. 166.

Modelling of sound transmissionthrough lightweight elements with stiffeners

M. Villot and C. Guigou-Carter CSTB, 24 rue Joseph Fourier, 38400 Saint-Martin-d'Hères, France

The sound reduction index of lightweight panels with stiffeners is calculated using the wave approach applied to an infinite thinplate line connected to periodically spaced beams and excited in flexure by a single or random (diffuse field) incident planeacoustical wave. The reaction forces and moments at the line connections are calculated from the flexural and torsional lineimpedance of the beams. The finite size of the panel is taken into account using a spatial windowing technique applied to theinfinite panel. The case of two stiffened panels with a cavity in between and no structural connections between panels is alsoconsidered using the same approach; using SEA, the structural flanking path at the panel boundaries through any receivingstructure can also be taken into account. Comparisons to measured results are given and discussed.

INTRODUCTIONIn order to understand the vibro-acoustical behavior oflightweight panels with stiffeners, a calculation modelbased on the wave approach has been developed. In thismodel, a thin plate is line connected to periodicallyspaced beams and excited by a diffuse acoustic field.The reaction forces and moments at the line connectionsof the stiffeners are calculated from the flexural andtorsional line impedance of the beams. The finite size ofthe panel is taken into account using a spatialwindowing technique applied to the infinite panel [1].Both the cases of a single panel and a double panelseparated by an absorbing cavity are considered. Fordouble panel systems, a structural flanking pathbetween the panels at their boundaries (through the testroom aperture or any other receiving structure) oftenexists. In this work, it is taken into account using a SEAapproach. Comparisons to measured results are givenand discussed.

(a) (b)

Stiffener

FIGURE 1. Schematic of investigated systems;(a) Single plate and (b) Double plate with stiffeners.

ANALYTICAL MODELThe 3-dimensional theoretical model is now shortlydescribed for a single plate with periodically spacedstiffeners as shown in Figure 1(a); it can easily beextended in the case of a double plate system asdepicted in Figure 1(b). The panel is excited by a

diffuse acoustic field composed of plane wavesimpinging at angles of incidence (θ,φ). The system isassumed to be surrounded by air (characterized by itswavenumber k0) on both sides. The time dependencehas the form ejωt (where ω is the angular frequency) andwill be omitted in the following for brevity. The normalvelocity of the panel is denoted v(y,z) in the spatialdomain and )k,k(v~ zy in the wavenumber domain. Thestiffeners are described by their spacing distance L,density ρr, Young modulus Er, shear modulus Gr,section area Sr, torsional constant Tr, flexural Ir andtorsional Γr moment of inertia. The reaction forces andmoments at the line connecting the stiffener to the plateare given by

)j(/)nLz()z,y(v)SkIE()z,y(F 2rr

4yrrr ω−δωρ−= (1)

)j(/)nLz()z,y(v)kTG()z,y(M 2rr

2yrrr ω−δ′ωΓρ−= (2)

where δ(z) is the delta Dirac function.The equation of motion giving the normal velocity ofthe plate is

)3()z,y(M)z,y(F

ekkk

cosk1P)z,y(v)(Z

n

nr

n

nr

)zcossinysin(sinjk

2z

2y

20

0inc

0

∑∑+∞=

−∞=

+∞=

−∞=

φθ+φθ−

−−

−−θ+=ω

where Z(ω) is the plate total wave impedance, includingthe mechanical and radiation impedance (on both sides).Taking the Fourier transform of equation (3) allowsobtaining an equation in the wavenumber domain, to besolved for each frequency ω and incident angle (θ,φ)

∑

∑+∞=

−∞=

+∞=

−∞=

π−

π−

ωωΓρ−

+

π−ω

ωρ−−

−φθδ−φθδ=ω

n

nzy

2

z

2rr

4yrr

n

nzy

2rr

4yrr

z0y0inc

zy

)Ln2k,k(v~

Ln2k

LjkTG

)4()Ln2k,k(v~

LjSkIE

)kcossink()ksinsink(P2)k,k(v~)(Z

)k,k(v~ zy is non zero for ky,n=k0sinθsinφ-(2nπ/L) andkz,n=k0sinθcosφ-(2nπ/L) with n integer between −∞ and+∞. To solve equation (4), the infinite sum(corresponding to the infinite number of stiffeners) is tobe truncated while insuring the convergence of the sum.Once the wavenumber spectrum is determined for eachfrequency ω and each angle of incidence (θ,φ), thetransmission loss for a diffuse incident field can then becalculated [2]. The finite size of the plate is taken intoaccount using the approach presented in [1].In the case of a double panel system, the transmissionloss associated with the vibrational flanking path at thepanel boundaries through the (usually heavier) receivingstructure, is taken into account using a SEA approach.For two identical panels, it can be shown to be given by

σρηπ

+= 2c

20

20

32

1023fl fcfm2

log10DTL (5)

where fc is the panel critical frequency, σ its radiationefficiency (see [3] for example), m its mass per unitarea, η its damping factor and ρ0c0 the air characteristicimpedance. D23 represents the vibration level differencebetween the two panels (measured according to [4]).

RESULTSFigure 2 shows the transmission loss for a 4 mm thickaluminum plate, without, and then with aluminumstiffeners (see [2] for description) periodically spacedwith a 40 cm distance. The finite size of the panel(3.43 m by 2.73 m) is considered. The calculated resultsagree well with the measured results : the stiffenershave no effect in the low frequency range; they lowerthe TL in the mid frequency range (below criticalfrequency) since the radiation efficiency is increasedand have a damping effect at the critical frequency.Figure 3 shows the transmission loss for a double platesystem used in buildings. Each plate is made of twogypsum boards (18+13 mm) screwed on the metalframe and the 20 cm thick cavity is filled with glasswool (σ=5 kPa s/m2). The stiffeners used (two U shapedsteel beams screwed together) are periodically spacedwith a 60 cm distance. The plate total loss factor η aswell as the vibration level difference D23 between theplates were measured. Figure 3 shows clearly the effectof the stiffeners (similar to that observed for the singleplate), starting above the 160 Hz third octave andlowering the TL by more than 10 dB around 1 kHz. Asa result, the flanking structural path between the panelswith stiffeners becomes dominant only at higherfrequencies; its effect can only be seen at and above thecritical frequency of the plate. Whereas, for the panelswithout stiffeners, the effect of the flanking structuralpath is of importance below and above the criticalfrequency.

0

10

20

30

40

100

125

160

200

250

315

400

500

630

800

1000

1250

1600

2000

2500

3150

4000

5000

6300

8000

Frequency (Hz)

TL (d

B)

Measured - Without Stiffeners

Calculated - Without Stiffeners

Measured - With Stiffeners

Calculated - With Stiffeners

FIGURE 2. Transmission loss for a single plate withstiffeners.

20

30

40

50

60

70

80

90

100

50 63 80 100

125

160

200

250

315

400

500

630

800

1000

1250

1600

2000

2500

3150

4000

5000

Frequency (Hz)

TL (d

B)

Measured - With Stiffeners

Calculated - Without Stiffeners - Without Flanking path

Calculated - With Stiffeners - Without Flanking path

Calculated - With Stiffeners - With Flanking path

Evaluated TLfl - Flanking path

FIGURE 3. Transmission loss for double stiffened panelsystem.

REFERENCES1. M. Villot, C. Guigou-Carter and L. Gagliardini, “Predictingthe acoustical radiation of finite size multi-layered structuresby applying spatial windowing on infinite structures,” Toappear in Journal of Sound and Vibration, (2001).2. F. Fahy, Sound and structural vibration – Radiation,transmission and response, Academic Press, 1985.3. R.J.M. Craik, Sound transmission through buildings usingstatistical energy analysis, Gower Publishing Ltd., 1996.4. CEN/TC126/WG6/N63, “Laboratory measurement of theflanking transmission of airborne and impact noise betweenadjoining rooms”, Part 1 : frame document.

Noise in Pipe Systems from Structure-borneEmission by Pumps

Q. Ning and B. Gibbs

Acoustics Research Unit, The University of Liverpool, PO BOX 147, Liverpool L69 3BX

In circulation systems, pumps deliver some of their acoustic power as structure-borne sound that propagates via the pipe walls asbending, axial and torsional waves. Wave mode conversion occurs at pipe junctions and it is not clear which component of thepower into the pipe system contributes most to the bending fields on system elements that radiate sound. A central heating systemwas modelled as various combinations of pipes and radiators. The input power was measured directly and the radiator energy wasobtained from the spatial average bending velocity. Results indicate that both the axial component and bending vibrationscontribute to the radiated sound. The experimental measurements were of fixed pipe lengths and resonance effects dominated thesystem response. In order to extract the general trends, a computer model of pipe systems was developed where the length of theconnecting pipes was varied stochastically. The theory demonstrates that at high frequencies the mixing of wave types is suchthat the radiated sound is determined by the largest power from the pump, irrespective of the component of excitation.Imperfections in the measurement set-ups resulted in more mode conversion and effective mixing occurred over the wholefrequency range of interest.

INTRODUCTION

Circulation pumps are the principal active componentin pipe systems and complaints of excessive noise indomestic central heating systems often result fromthem. Problems are seldom due to pumps radiatingsound directly into the air. The noise results from fluid-borne sound transmission through to the connectingpipe and radiator systems and from structure-bornesound transmission that directly excites the pipe walls,radiators and the supporting walls and floors.Therefore, the problem is determined in part by theinstallation as well as the pump.

It has been demonstrated that pumps can be consideredas sources of structure-borne and fluid-borne soundsimultaneously, by considering emission into semi-infinite pipe systems [1]. The fluid-borne emission canbe measured directly in a test rig of flexible pipeswhere the returning propagating waves are muchattenuated. The structure-borne emission cannot bemeasured directly but is obtained from measurement ofthe pump free velocity sfv and mobility sY , and from

the predicted mobilty rY of a semi-infinite pipe ofmaterial and cross-section typical of domestic heatingsystems. The active power transmitted then is obtainedfrom:

)Re(||

||21

2

2

rrs

sf YYY

vP

�

�

The structure-borne sound propagation along the pipesis predominately in the form of axial and bending

vibrations which partially convert to each other at pipejunctions [2]. It therefore is not immediately obviouswhich component of the input power contributes mostto the radiated sound, except in very simple systems.However, it can be postulated that complicated pipesystems may ‘mix’ the contributions of the inputcomponents efficiently so that the magnitude of theinput power is the determining factor irrespective ofthe direction of excitation. An experimental andtheoretical investigation of wave mode conversion hasbeen conducted in order to determine the relativecontributions of the structure-borne emissions frompumps to the noise radiated from pipe systems.

EXPERIMENTAL INVESTIGATION

A series of pipe-radiator systems of increasingcomplexity were assembled. The power input at thefree end of a pipe was measured directly from the crossspectrum of the applied force and velocity. Theradiator energy was obtained as a spatial average of thebending field. In Figure 1 is shown the ratio of thebending energy of a radiator, connected to a three-pipesystem, to the input power, for three mutuallyperpendicular directions.

The relative contributions are more clearly seen inFigure 2 where the two components, perpendicular tothe input pipe axis are normalised with respect to theaxial component.

0.0000001

0.000001

0.00001

0.0001

0.001

0.01

0.1

1

10 100 1000 10000Frequency, Hz

Tra

nsm

issi

on c

oeff

icie

nt

Fig 1. Measured ratio of bending energy of a heatingradiator to the three translational components of power inputto a three-pipe system.

0.0001

0.001

0.01

0.1

1

10

100

1000

10 100 1000 10000

Frequency, Hz

Pow

er r

atio

Fig 2. Measured ratio normalised with respect to the axialcomponent of input power.

Results indicate that both the axial component andbending vibrations contribute to the bending fields ofthe radiators and none can be neglected a priori.Indeed, the three components assume equalimportance, even with relatively simple systems ofthree pipes and two junctions. The experimentalmeasurements were of systems of fixed pipe lengthsand resonance effects dominated the system responseand therefore obscured trends in wave conversion.

NUMERICAL MODEL

In order to extract general trends from systemresonance effects, a computer model of pipe systemswas developed where the length of the connectingpipes could be varied stochastically in order to extractthe trends due to mode conversion. Figure 3 shows thepredicted ratio of output power to that input for threemutually perpendicular pipes. The three ratiosconverge with increased frequency.

-20

-18

-16

-14

-12

-10

-8

-6

-4

10 100 1000 10000

Frequency, Hz

Tra

nsm

issi

on c

oeff

icie

nts,

dB, r

ef. =

1.

Longitudinal input Bending input y-direction Bending input z-direction

Fig 3. Ratio of bending power on a receiver pipe to that on asource pipe, for a three-pipe system.

CONCLUDING REMARKS

It has been demonstrated experimentally that pipe-systems with more than two junctions, effectively mixthe contributions from the various components ofexcitation by the pump. A computer model alsopredicts energy equi-partitioning but at higherfrequencies. This is to be expected since there are no‘imperfections’ in the computer model. In realinstallations, there will be imperfections in theassembly which also contribute to mode conversion.Then, the component of pump excitation of greatestmagnitude is the dominant contribution to the radiatedsound, irrespective of the direction of excitation.

REFERENCES1. B.M.Gibbs and Qi Ning, “Fluid-borne and Structure-borne

Sound Emission from Small Circulation Pumps”, Proc. 7th

International Congr. on Sound and Vibration, 1597-1604, 2000

2. Qi Ning, “Towards a Characterisation of Pumps as Sources ofStructure-borne Sound”, Ph.D. Thesis, Liverpool University,2000.

Solution of the acoustical Problems for the membraneroofof the NBIA New Bangkok International Airport

A. Egentera,b, D. Schräpela,c, R. Bluma, H. Bögnera

Laboratorium Blum, Handwerkstrasse 58, 70565 Stuttgart, Germany, Labor_Blum@compuserve.comClear-Sound-Acoustics, Hauptstrasse 119, 73312 Geislingen, Germany, CLEAR-SOUND-STUDIO@t-online.de

Universität Stuttgart, Pfaffenwaldring, 70569 Stuttgart, Germany

The roof of the concurses of the NBIA is covered besides glass by a PTFE-covered glass fibre fabric. Both, soundinsulationand roomacoustics of this material can only be judged as extremely bad. Through the application of a further developedacousticelement both, the soundinsulation against noise by aircraft and the roomacoustics could be improved essentially. Thiscould be reached by a multilayer-structure. The translucence and the low mass per area of the roof, both essential propertiesof membrane-buildings, could be maintained. The surface has to have Low-e-properties, for the reason of energy theemissivity in the infrared range has to be low. This requirement could also be kept with this structure.

SOURCES OF NOISE

For the noise pollution level of an at least 12 year oldB has been used. The noise of an old Boeing 707 is107 dB(A) in a distance of 400 metres. Besides thealready dealed takeoff noise, there is a second noisesource, the noise of the aero-engines at the drive totheir end position.

CONCEPT OF NOISE REDUCING

For effective noise-reducing the roof has to have arelatively good sound insulation and it has to controllroomacoustics. Where the sound-pressure-level ishigh (see Figure 1). If the membrane oscillation ofthe outer and inner membrane are in phaseopposition, the air between the membranes will becompressed.

FIGURE 1. Principe of the relation between membraneoscillation and the distribution of pressure between themembrans

For this reason, a sound-pressure between themembranes comes about. The maximum of thissound-pressure is in the middle between themembranes. In this way, a part of sound-energy willbe taken away from the sound-field.

Sound insulating Layer

The mass of membranes is very small for soundinsulation. The sound-insulation can be improved bya mass-layer between the two membranes. Thismass-layer has sound-reflecting properties assketched in Figure 2.

FIGURE 2. Three-layer-system without absorbingelements

So, a part of the sound will be reflected. For thatreason resonances between membrane and mass-layer come in beginning. By these resonances thesound insulation gets worse.

Combined sound-insulating andabsorbing intermediate layer

incidentNoise

Outermembrane

Insulatinglayer

InnermembraneMembraneoscillation

Distribution of pressurebetween the membrans

Out

er m

embr

ane

Abs

orbe

r .

insi

de m

embr

ane

To prevent these resonances, an absorbing layer canbe mounted at both sides of the sound-insulatingintermediate layer as shown in Figure 3. This sound-energy is not disposable at the inner membrane. Forthis reason these resonances will be suppressed, andthe sound insulating of the whole structure willincrease.

FIGURE 3. Three-layer-system with absorbing elements

Membrane-buildings are leight-weighted and trans-lucent. For this reason we have developed a newacoustic element for membrane-buildings (so calledbaffle) which combines all the requirements for thesecase as: leight-weighted, translucent, sound-insulating, sound-absorbing and UV-resistant.

INTERMEDIATE LAYER (BAFFLES)

The physical principles on what the function of theintermediate layer, so called baffles bases are thestimulation of bending waves and compression of air[1.]. In figure 4 the absorption-coefficient of thebaffles are shown for a baffle which is directmounted on a wall, and at a distance of 200 mm.

FIGURE 4. absorbing coefficient of the baffles

The sound-insulation of the baffles is shown inFigure 5.

FIGURE 5. sound-insulation of the baffles

For room-acoustics, the effect of the baffle dependsof the transmission of the inner membrane. Acomparison of the sound-insulation of the membraneand the sound-insulation of a system of baffle andmembrane is shown in figure 6.

FIGURE 6. sound-insulation of the Membrane with andwithout baffle

The combination of baffle and membrane has clearlyhigher sound-insulating-values, than the membranealone. And this by a minimal addition of the totalmass of the whole system.

RESULTS

The sound reduction only of the membrane is verybad - as expected. The additional baffle layer resultsin better values. The sound transmissioncharacteristic Rw' comes out to be 26 dB for onelayer of baffles in the distance of 40 cm to themembrane and 30 dB for the two layers. So the noiseof an old B747 will be reduced to approximately 76dB(A) in the worst case: In the part of the concoursesnext to the runways.. The noise of an A300 will bereduced to a value lower than 70 dB(A) in allsituations.

CONCLUSION

If the properties of the membranes like mass persquaremetre, distance between the membranes andsound-absorbing as well as sound-insulatingproperties of the baffles tunes to each other, there isan acoustic element which fulfills the requirements.

REFERENCES

1. Mechel, F. P., Kiesewetter, N.: Schallabsorber ausKunstoff-Folie, Acustica 47, 83 – 88, (1981)

0

5

10

15

20

25

30

35

40

2 2,3 2,6 2,9 3,2 3,5 3,8

frequency [Hz]

sou

nd

-in

sula

tio

n [

dB

]

100 200 400 800 1600 3150

membrane with baffel

Baffel-distance 200 mm

Baffel-distance 40cm

membrane without baffel

Incident Noice

Outer membrane

AbsorberInsulating layer

Absorber

Inner membrane

0,00

5,00

10,00

15,00

20,00

25,00

30,00

35,00

40,00

45,00

100 1000 10000

frequency [Hz]

0

0,2

0,4

0,6

0,8

1

1,2

2 2,3 2,6 2,9 3,2 3,5

frequency [Hz]

abso

rpti

on

co

effi

cien

ta

100 200 400 800 1600 3150

Air-layer 200 mmdirect mounted

Comparison between laboratory and in situ Sound Reduction Index measurements

S. Secchia and P. Faustib

aDepartment of Processes and Methods of Building Production, University of Florence, Italy

bDepartment of Engineering, University of Ferrara, Italy

The paper deals with the results of research concerning the acoustical properties of masonry walls tested both in a laboratory with suppressed flanking transmission and in a laboratory with relevant flanking transmission (in situ). Measurements were aimed at evaluating the difference between Sound Reduction Index in laboratory and in situ and the amount of flanking transmission through different lateral paths. The results confirm the reliability of the flanking transmission calculation model defined by the European Norm and give a contribution to the set-up of the new method of measurement of the vibration reduction index.

INTRODUCTION

Sound transmission between adjoining rooms is due to energy propagation both through the separating wall and through the lateral structures of the two rooms. The European Norm 12354-1 [1] makes it possible to estimate the acoustic properties of buildings through the calculation of both these forms of transmission. The paper presents the results of research concerning the acoustical properties of masonry walls tested both in a laboratory with suppressed flanking transmission and in a laboratory with relevant flanking transmission (reproducing a typical real building context). Flanking transmission has been evaluated by measuring the vibration reduction index Kij, whose test method is studied in ISO/CD 10848 [2], and the structural reverberation time Ts of the partition wall and of the lateral structures. DESCRIPTION OF THE TEST WALLS AND MEASUREMENT PROCEDURE

Five different masonry walls were tested in two laboratories with and without flanking transmission. The laboratory with suppressed flanking transmission respects the recommendations of ISO 140-1 [3], while the other one respects the recommendations of ISO/CD 10848 [2]. The realisation of the walls was carried out in one laboratory by a single firm with a single supervisor and with strictly defined construction procedures. The same thing applied to the other laboratory. Four walls were single layer while the other one (E with reference to table 1) was double layer. A synthetic description is given in table 1. The sound reduction index was measured in laboratory according to ISO 140 – 3 recommendations and in situ according to ISO 140 – 4.

Table 1. Synthetic description and rating of thesound reduction index of the five walls tested. Wall A B C D E Surface mass (kg/m2) 470 370 400 390 320

Thickness (m) 0,45 0,28 0,21

0,33 0,17 0,03 0,08

Rw (measured in laboratory) 50 53 54 56 52 R’w (measured in situ) 47 49 48 50 47 R’w (estimated in situ) 45 46 46 47 45

The vibration reduction index, Kij, is the new quantity introduced by EN 12354-1 to evaluate the flanking transmission and can be calculated by means of simplified formulas given by EN 12354-1 [1] or measured in situ, according to the recommendations of ISO/CD 10848 [2]. For the estimation, the kind of junction between the test wall and its lateral structures needs to be known. In the research, where all the junctions were rigid, Kij was both estimated and measured. The sound reduction index of all the lateral structures of the two adjoining rooms were known from previous laboratory measurements. From the measured values of the sound reduction index of the test wall and of the lateral structures and from the measured vibration reduction indexes of all flanking paths, the apparent sound reduction index R’ was calculated. The results obtained were finally compared with the measured values of R’.

RESULTS

Table 1 shows the comparison between the estimated and the measured values of R’w. It is evident that in the case studied, the estimation model of EN 12354-1, with the measured values of Kij, gives lower values of R’w than in real life. This is due to the presence of some lateral paths which are characterised by very small values of the vibration reduction index and, as a

consequence, by very small values of Rij. In particular, paths 2-6 and 4-8, concerning the vibration transmission from the two floors of the source room to the corresponding floors of the receiving room, are characterized by much smaller measured values of Kij than the estimated values. Figure 1 shows, for two walls, the comparison between the estimated (grey) and the measured (white) values of Kij (odd numbers indicate the lateral walls, while even numbers the lateral floors). Figure 2 shows, for the same two walls, the comparison between the measured and the estimated values of R’ and the measured value of R, as a function of the frequency.

-4

-2

0

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16

K 3-S K S-7 K 3-7 K 1-S K S-5 K 1-5 K 2-S K S-6 K 2-6 K 4-S K S-8 K 4-8

Frequenza (Hz)

Vib

ratio

n re

duct

ion

inde

x dB

)

measured estimated

-6

-4

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0

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K 3-S K S-7 K 3-7 K 1-S K S-5 K 1-5 K 2-S K S-6 K 2-6 K 4-S K S-8 K 4-8

Frequenza (Hz)

Vib

ratio

n re

duct

ion

inde

x dB

)

measured estimated FIGURE 1. Comparison between estimated and measured values of Kij (mean values between 500 and 2000 Hz) (wall C above and wall E below).

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65

100 125 160 200 250 315 400 500 630 800 1000 1250 1600 2000 2500 3150

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)

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100 125 160 200 250 315 400 500 630 800 1000 1250 1600 2000 2500 3150

Frequency (Hz)

Sou

nd r

edu

ctio

n in

dex

(dB

)

R (laboratory) R' (theoretic) R' (situ)

FIGURE 2. Comparison between values of R (measured) and R’ (measured and estimated) (wall C above and wall E below).

CONCLUSIONS

The comparison between measured and estimated values of the apparent sound reduction index (figure 2) shows a close correspondence if the differences between the realisation of the walls in laboratory and in situ are taken into account. The estimation model of EN 12354-1, using the measured values of the vibration reduction index, can explain the reduction in the isolation from the laboratory to real life.

ACKNOWLEDGEMENTS

The research, carried out by the universities of Ferrara (supervisor), Padova (for the laboratory measurements according to ISO 140-3) and Trento (for the in situ measurements according to ISO 140-4), was founded by Italian Association of Brick Industrialists (ANDIL) and by National Research Council (CNR), project nr. PF MSTAII n°99.01848.PF34.

REFERENCES

1. EN 12354-1 Building Acoustics - Estimation of acoustic performance of buildings from the performance of products, part 1, Airborne sound insulation between rooms.

2. ISO/CD 10848, Acoustics - Laboratory measurement of the flanking transmission of airborne and impact noise between adjoining rooms.

3. ISO 140, Acoustics - Measurements of sound insulation in buildings and of buildings elements.

Improvement of sound insulation of building facades

H. Tachibanaa and S. Sakamotob

a, b Institute of Industrial Science, University of Tokyo, Komaba 4-6-1, Meguro-ku, Tokyo 153-0041, Japan

For the mitigation of road traffic noise problem, it is needed to improve the sound insulation of facades of roadside buildings as well as the reduction of noise generation from road vehicles and acoustical improvement of road structures. In this paper, the effects of sound absorption treatment for balcony space of roadside buildings are investigated by scale model experiment, numerical analysis and field measurement.

INTRODUCTION According to the result of the field measurements of road traffic noise in Tokyo performed in 1999, the attainment ratio of the “Environmental Standards for Noise” [1] is 30 percent for daytime and only 12 percent for nighttime. In a lot of areas, the noise level exceeds 75 dB. In the “Environmental Standards” revised in 1999, it is specified that the noise level inside the residential buildings should be less than 40 dB in LAeq in nighttime in the areas facing artery roads. To preserve this value, roadside buildings must have sound insulation of more than 30 or 35 dB. It is rather difficult to realize such high sound insulation by the ordinary façade constructions and therefore various kinds of devices are needed to improve the sound insulation of buildings. As a study of such kind of building acoustic research, we have made experimental and numerical investigations into the effect of sound absorption treatment for the balcony space facing to a road.

THE EFFECT OF SOUND ABSORPTION TREATMENT FOR BALCONY SPACE

(1) Study by 1/16 scale model experiment As the first study, 1/16 scale model experiment was performed. In this experiment, a chest of drawers was used as the model building (see Fig.1). In a drawer, a model room with a window of 25 cm x 12.5 cm (4 m x 2 m in full scale) made of 0.1 mm thick aluminum plate was made and it was moved to each floor. The sound pressure level in the model room was measured under the conditions of with and without sound absorption treatment for the ceiling of the balcony. Wool felt of 2 mm thick was used as the sound absorber. From the level difference between these two conditions, the effect of the treatment was examined. As a result, it has been found that the sound insulation increases up to 5 dB on the floors upper than the first floor by the absorption treatment as shown in Fig.1.

(2) Study by numerical simulation As the second study, numerical simulation using the finite difference time domain (FDTD) method [2] was performed. In this study, the propagation of an impulse (a half period of a sine wave) generated from a sound source in 2-dimensional space was calculated. An example of the calculation result is shown in Fig.2 which shows the instantaneous wave forms after 56 ms. By calculating the difference of squared and integrated impulse response on the surface of the window between the conditions of with and without sound absorption treatment, the effect of sound absorption was examined. In this case, the window surface was assumed to be acoustically rigid, and 0.7 sound absorption coefficient was assumed for the absorption treatment. The calculation results are shown in Fig.2, in which the tendency almost similar to the result of the scale model experiment is observed. This result is almost the same as that of the numerical study made by D.C. Hothersall et al. [3]. (3) Study by field measurement In the previous two studies, it has been found that the sound absorption treatment for the ceiling of the balcony space is effective for the improvement of sound insulation of building façade. To confirm this effect in an actual building, field measurement was performed in a residential building facing a railway (see Fig.3). In this study, the mean sound pressure level in a room on the third floor and the sound pressure level just outside the building during the passage of trains were measured simultaneously, and the level difference between inside and outside was obtained. This measurement was made before and after the sound absorption treatment for the ceiling of the balcony. From the change of the sound pressure level difference before and after the sound absorption treatment, the extent of sound insulation improvement was obtained. As a result, it has been found that the sound insulation has been improved by 3 to 7 dB in the frequency range of 125 Hz to 2k Hz by the sound absorption treatment as shown in Fig.3.

CONCLUSIONS From the results of experimental and numerical investigations mentioned above, it has been confirmed that the sound insulation of building façade can be improved by performing sound absorption treatment for the ceiling of the balcony space. Beside the acoustical treatment mentioned here, other devices are further needed for the improvement of sound insulation of buildings locating near such noise sources as artery roads and railways. In the future, the possibility of double-skin construction with ample air space should be considered to realize sufficient sound insulation performance.

REFERENCES 1) NNI Editorial Staff, “Overview of Japanese

Environmental Regulation on Noise,” Noise/News International Vol.8, No.2, 64-76 (2000.6)

2) S. Sakamoto, Y. Tokita and H. Tachibana, “Calculation of impulse responses of rooms by using of the finite difference method,” Proc. Of ASA & ASJ Third Joint Meeting, 1307-1310 (1996)

3) D.C. Hothersall, K.V. Horoshenkov and S.E. Mercy, “Numerical modeling of the sound field near a tall building with balconies near a road,” J. Sound & Vibration, 198(4), 507-515 (1996)

Fig.1 Study by 1/16 scale model experiment

216

cm

104 cm

Sound source (horn tweeter)

Mic.

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tion

incr

ease

[dB

]

125 250 500 1k 2k-5

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15

Frequency [Hz]

3F

1F 4F 2F 5F

Fig.3 Study by field measurement

Sound absorption treatment

20.8

m

1.6m 11.6m 3.5m

Mic.

Train

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tion

incr

ease

[dB

]

63 125 250 500 1k 2k -5

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15

Frequency [Hz]

Fig.2 Study by numerical simulation (FDM)

2.9

m

S

8 m 2.82 m 1.6 m

R

Absorption

(a) without absorber (b) with absorber

Soun

d in

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tion

incr

ease

[dB

]

125 250 500 1k-5

0

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15

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3

1 42 5

6

Flanking Sound Transmission along an AluminiumFaçade - experiments vs Prediction -

H.J. Martin, M.E.A. Schoffelen and W.M. Siebesma

Department of Architecture, Building and Planning, Eindhoven University, PO. Box 513,5600 MB Eindhoven, The Netherlands; h.j.martin@bwk.tue.nl

In an apartment building aluminium facades are applied in front of the concrete floor between the apartments. This results in avery complex junction. The airborne and impact sound insulation between the apartments should fulfil comfort class demands:Rw ≥ 57 dB and LnTw ≤ 50 dB. A prediction of the airborne sound insulation was made based on the proposed building elements.The (newly developped) façade is not under construction yet. In the laboratory test have been done on a T-junction consisting ofa concrete floor and a similar, substitute façade. The measured airborne sound insulation just fulfilled the demands. Also thevibration transmission indices Kij between all elements of the T-junction have been measured; the measured values have beenused as input for the prediction model according to EN 12354-1. Comparison of the measured and calculated data show that thereis quite a good agreement when the measured values of Kij are used as input data for the junction characteristics, and largedifferences will occur depending on the way the junction is or can be modelled.

INTRODUCTION

Aluminium facades are mainly applied in officebuildings. Applying these types of facades in high-riseapartment buildings with higher acoustical demands,flanking transmission along the facade may becomeimportant and maybe dominant, especially when thefacade is situated in front of the (concrete) floorbetween the apartments and when a higher comfortclass is aimed at.These problems were met in a real case, underdevelopment. Here acoustical demands betweendwellings were higher than building regulationsminimum values: Rw ≥ 57 dB and LnTw ≤ 50 dB.The demands for impact sound insulation are metrather easily using a heavy floating floor construction.A prediction of the airborne sound insulation was madebased on the proposed building elements.Questions about the accuracy of the prediction arose,so tests have been done in the laboratory on the T-junction of floor and facade to determine the overallsound insulation as well as the vibration tranmissionacross the junction. By using these results in a newprediction we tried to confirm the first predictionresults. Only a substitute facade had to be used in thetest because the real facade was not under constructionat that moment. Only the direct sound transmissionthrough the concrete floor and flanking transmissionvia the facade have been taken into account in thelaboratory tests.

JUNCTION ELEMENTS

The T-junction of floor and facade consists of a 220mm concrete floor with a 70 mm floating floor resting

on 20 mm rockwool and an aluminium facadecontaining two types of double glazing: 8-15-6 mmnormal thermal insulating glass and PVB-laminatedsafety glass 8-12-44.2 mm.The relevant part of the facade consists of threeelements: on each concrete floor an aluminium facadeelement with glass is mounted hanging; in front of eachconcrete floor a coupling element is mounted betweentwo facade elements. This coupling element is madeof aluminium with plywood and filled with rockwool.It is connected to the upper and lower facade elementsby means of flexible rubber profiles.This results in a very complex junction.

TEST SETUP

The above described junction has been mountedbetween two transmission rooms with the followingdifferences:- no floating floor present;- substitute facade elements have been used having acomparable thickness, stiffness and surface mass.Dimensions of the junction:- 210 mm concrete floor, area 6,7 m2;- upper facade area 9,2 m2;- lower facade element area 10,9 m2;- coupling length between floor and facade 3,3 m;- each facade element contains three verticalaluminium studs; the upper element has threehorizontal studs, the lower element has four horizontalstuds;

- sound transmission along other paths is minimized bymeans of double metal-stud walls.

- source and receiving room volumes respectively 59,8and 73,7 m3.

MEASUREMENTS

The airborne sound insulation between source andreceiving room has been determined according ISO140-3.The vibration transmission indices Kij between theconcrete floor and both facade elements have beendetermined by exciting each element successively andmeasuring the vibration levels and the structuralreverberation time on both relevant elements.The facade was excited by a vibration exciter, theconcrete floor by the ISO tapping machine.

RESULTS

Overall sound insulation

Overall sound insulation just fulfilled the demands: Rw= 58 dB. Because of the absence of the floating floor,an even higher sound insulation may be expected inpractice, assuming sound transmission via other pathscan be neglected.

Junction characteristics

Figure 1 shows the measured vibration transmissionindices between floor and both facade elements inthird-octave bands. These values have been used asinput data for the prediction of the airborne soundinsulation between the transmission rooms.

PREDICTION RESULTS

The over-all sound insulation has been calculated witha computer model based on EN 12354-1. There aredifferent ways to model this complex junction into thecomputer model. In the facade elements, the doubleglazing is the dominant surface, so it seems logic tomodel the facade elements by means of the glazing as alightweight flanking construction. However, theinfluence of the flexible couplings is hard to estimate.Therefore, the measured values of Kij have been usedas input data for another calculation.Comparing the results of prediction and measurementsit seems that:- about 90 % of the sound is transmitted via the directpath; - when the measured values of Kij are used asinput data, the differences between prediction andmeasurement results are about 1 dB.- when the facade elements are modelled as doubleglazing, the calculated results stay behind the measuredones by about 8 dB;

0

10

20

30

40

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63 125 250 500 1000 2000 4000

frequency [Hz]

Kij

[dB

]

between facades lower façade-floorupper façade-floor

FIGURE 1 Vibration transmission indices of the junctionfacade-floor

CONCLUSIONS

The combined results of prediction and laboratory testconfirmed the first prediction results, so it seems thatthe acoustical demands can be met in practice.The best way to model this type of facade in flankingtransmission calculations seems to be as double glazingwith an extra 5 dB in Kij to bring in the effect of eachflexible connection.

REFERENCES

EN 12354-1: Building acoustics – Estimation ofacoustic performance of buildings from theperformance of elements –Part 1: Airborne soundinsulation between rooms

An Invariant to Predict Noise Insulation of Beam-Block FloorsJ. Patrícioa; L. Bragançab

a - LNEC, Av. do Brasil, 1700 Lisboa, Portugal; e-mail: jpatricio@lnec.ptb - UM, Dep. Eng. Civil, Azurém, 4800 Guimarães, Portugal; e-mail: braganca@eng.uminho.pt

There has been a huge abandonment of inhabitants from most part of portuguese city centres. The latest governmental and socialpolitics go towards specific developments to bring back people. So, the existing antique buildings need to be rehabilitated,converted, reinforced, etc., and, consequently, the acoustic insulation will be an important aspect to take into account. In Portugal,for rehabilitation purposes, the beam-blockt floors are the most commonly used. Nevertheless, in what respects their acousticperformance there are some doubts from householders and users concerning the existence of non-conformities, in terms of impactnoise insulation with national regulations. This comes straightway from the fact that the simplified acoustic models used to predictimpact noise insulation are not as suitable as they should be for the case. This paper presents some considerations about this aspectand proposes a practical model based on the results of a set of tests carried out in situ and in laboratory which has led to an invariant(Rw + Ln,w) for these type of horizontal building partitions.

INTRODUCTION

Nowadays and based on the need to lighten thepermanent loads in buildings as well as to adopt newtechnological constructive solutions, it is widely usedlightweight floors. In Portugal, those most used,particularly in rehabilitation processes, are the beam-block, as figure 1 illustrates.

Fig. 1 – Typical beam and pot floor

Regarding impact noise and for acoustical designingpurposes, these floors are seldom considered massive �3�. Despite the existence of several potential theoreticalmethodologies to predict noise insulation of buildings,some doubts still remain. This paper shows that thesimplified model often used by building agents:designers, authorities, etc., to predict noise insulation ofthese floors, against impact noise, does not work alwayswell. Thus, it presents some considerations about theacoustic performance of beam-block floors and proposesa practical model based on the results of a set of testscarried out in situ and in laboratory which has led to aninvariant (Rw + Ln,w) for these building partitions.

Theory

For reasons of reciprocity the sum of the airborne noisereduction index R and the normalized impact noisepressure level Ln for homogenous floor constructionsdepends only on frequency, if forced transmission isnegligible �1�. In this situation the impact noise pressurelevel of the construction can be estimated from data onthe noise reduction index and vice-versa, according to thefollowing equation in third octave bands of centrefrequency f:

(1) dB ]Hz1[

flog3038LR n ���

Generally, the estimation of impact noise insulation isnot so well known as the airborne noise insulation offloors is, because of the mass law. Considering the forcelevel of the standard tapping machine LF according to ENISO 140-6, we will have for one third-octave bands theequation:

� � � �

(2) dB fflog10log10

s1T

log10m/kg 1

mlog30155L

.ref

s22n

���

��

��

where TS is the structural reverberation time; � theradiation factor of the floor for free bending waves andfref. is the reference frequency (fref. = 1 kHz).

However for homogenous floor costructions theequivalent weighted normalized impact sound pressurelevel Ln,w,eq used to determine, in a simplified way, thecalculation of impact noise insulation of floors can becalculated from the mass per unit area m´, in the rangefrom 100 kg/m2 to 600 kg/ m2, with the equation:

� �(3) dB

kg/m 1m35-164LLL

2ww,n0,eq,w,n

�����

Considering that the beam-block floors can be assumedhomogenous in what concerns impact and airborne noise,it follows for the invariant in third-octave bands theequation (in this case �Lw = 0 dB):

(4) dB .InvLR 0,eq,w,nw ��

Results

These sets of floors were tested in laboratoryconditions. Form these results a proposal for an invariantwas envisaged. The corresponding values are in Table 1.

Table 1 – Values of the Invariante Ln,w + Rw for the beam andfloors tested

Floor Sound insulation indices

Ln,w(dB/oit)

Rw(dB)

Invariant(dB/oit)

F1 94,7 50,9 146F2 97,2 50,0 147F4 96,7 50,6 147F4 97,4 50,8 148P5 94,5 53,5 148

Based on these results and having in attention that theycome from 9 tests made on each floor (one impactposition for 3 receiving points in the room), the averagevalue obtained for this Invariant is 147 dB/oit with astandard deviation of 1 dB/oit. This Invariant seems tobe well correlated with values obtained in testsperformed in several current housing buildings, as Table2 illustrates. The buildings used are of current type builtin Portugal and the rooms are of typical dimensions inhousing buildings. The values obtained for both theweighted airborne noise and the weighted impact noiseinsulation are summarized in the Table 2, as well as therespective invariant and the values obtained with the

symplified prediction method set up in the europeanstandard �1�, Eq (4).

Table 2 – Values of the Invariante Ln,w + Rw for the double(DB) and single (SB) beam and pot floors tested on site

Site Rw(dB)

Ln,w(dB/oct.) Inv.

Ln,w,eq,0

(dB/oct.)

Floortype

AV1AV2AV3AV4AV5PB1PB2PB3IL1IL2

51494751525455515149

89939696949594949393

140142143147146149149145144142

77797977777777797977

DBSBSBDBDBDBDBSBSBDB

CONCLUSIONS

Based on the results obtained with these set of tests it isevident that the symplified method established in the neweuropean standard does not give suitable resultsregarding the impact sound insulation of bare floors. Thevalue of 147 dB/oit (142 dB) seems more appropriate.This is coherent with the average value for the invariantobtained in Table 2 (145 dB/oit. - that makes 140 dB). Concerning the effects of flanking transmission it ispossible to conclude that they are not so important as toturn the invariant unadequate. This conclusion comesfrom the fact that the flanking transmission for commonportuguese constructions - mansonry walls with mass perunit area varying from 120 kg/m2 to 150 kg/m2 –connected with X and T junctions with these floors - onlydecreases 3-4 dB the airborne noise insulation index and1-2 dB the impact sound insulation index, yielding acumulative difference of 2 dB for the invariant.

REFERENCES

[1] CEN– Building Acoustics. Estimation of acoustic performance ofbuildings from the performance of elements. Part 2: Airborne soundinsulation between rooms. EN 12354-1,2, 1999.[2] EN ISO 140-6,7: 1998 – Acoustics. Measurement of soundinsulation in building elements.[3] EN ISO 717-1,2: 1996 – Acoustics. Rating of sound insulation inbuildings and of building elements.[4] PATRÍCIO, J. V. – Acoustic performance of non-homogeneousfloors regarding impact sound in buildings: simulation model “Ph. D.Thesis”). LNEC, Lisbon, 1999.

Use of Finite Element Method to Investigate the Effect ofFurniture, Wall Recesses and Construction Materials on the

Sound Field in Dwellings at Low FrequenciesS. Maluski and B. Gibbs

Acoustics Research Unit, The University of Liverpool, PO BOX 147, Liverpool L69 3BX

In general the sound insulation at low frequencies in dwellings is poor and there are not yet accurate methods of laboratory andfield measurement of sound insulation at such frequencies. A Finite Element Method has been used to investigate the soundtransmission between rooms below 200Hz. By modelling rectangular rooms, it was observed that the room volume stronglyinfluences the sound pressure level difference of party walls. However, the FE models were not of real dwellings. Whencomplaints from residents occur, the rooms are fully furnished. Moreover, many rooms are not perfectly rectangular and mayhave recessed walls. In addition, dwellings are built with different types of material. A FEM model has been developed toinvestigate the effect of these three different parameters: furniture, wall recess and construction materials. Measurement in-situshowed that the effects of furniture are insignificant below 100Hz. Modelling a wall recess improved the agreement betweenprediction and measurements but the assumption of a pure rectangular room remains appropriate. A damping, equivalent to anabsorption coefficient of 0.02, reproduces the effect of masonry and a damping, equivalent to an absorption coefficient of 0.1,recreates the effect of plastered timber-frame walls, floors and ceilings.

1. INTRODUCTION

Building elements generally give poor sound insulationbetween dwellings at low frequencies. This is a majorcontributory factor to the increasing number ofcomplaints resulting from enhanced bass sounds fromhi-fi or home cinema systems, and from mechanicalservices and road traffic1,2. As low frequencymeasurements produce poor measurement repeatabilityand reproducibility, a Finite Element (FE) model wasused to investigate the sound transmission betweenadjacent rooms3. However, the model, althoughvalidated by measurements, was a simplification ofreal situations. It did not include the effect of roomwall and floor construction (ie. lightweight cavity orheavyweight masonry) or the effect of furnishing.Moreover, most rooms have wall recesses and cannotbe assumed perfectly rectangular.

This paper presents an investigation of the effects ofconstruction material, furniture and wall recess on theroom response at low frequencies. Field measurementdata is compared with FE prediction, to ascertain theseeffects and how best they can be modelled.

2. FIELD MEASUREMENTS AND FE MODELThe sound fields in living rooms and bedrooms weremeasured in dwellings of heavyweight masonry orlightweight cavity construction. The acoustic fields weremeasured at a single microphone position with a largespeaker, located in a corner. The frequency response of theroom was measured from 25Hz up to 205Hz. FE models ofthe rooms also were constructed for comparison.

3. CONSTRUCTION MATERIALS

Fig. 1 shows the measured and predicted frequencyresponse of a 5.75x4.88x4.24m room of plastered brickwalls and concrete floor and ceiling.

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No damping Surface absorption 0.02Measurement Eigenfrequencies

Fig 1. Measured and predicted frequency response of a roomof plastered brick walls and concrete floor and ceiling.

FEM prediction gave the best agreement withmeasurement if a surface absorption coefficient of 0.02was assigned.

Fig. 2 shows the measured and predicted frequencyresponse of a 4.24x2.84x2.40m room with plasterboardand timber-frame walls, floor and ceiling for a range ofsurface absorption coefficients. The predicted curvesdisplay the same signature as the measured. It showsthat a surface of assumed absorption coefficient of 0.15gives a best agreement with measurement.

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No damping 0.03 0.150.2 Measurement Eigenfrequencies

Fig.2. Surface absorption coefficient to model the frequencyresponse of a room with timber-frame walls, floor andceiling.

4. FURNITURE

The effect of four different sofa and chairs’ positionson the sound field of a living room was investigated; asshown in Fig.3.

A

DC

B

Fig.3. Four distributions of sofa and chairs

Fig. 4 shows that the introduction of furniture changesthe frequency response by 1-2dB below 100Hz, and by5dB above 100Hz. It can be assumed that furniture hasno influence on the sound field at such lowfrequencies.

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Fig.4. Level difference between unfurnished and furnished

It follows that the position of furniture has a negligibleeffect and this also is confirmed in Fig.4.

5. WALL RECESS

Many rooms in dwellings have wall recesses e.g.chimney breasts and alcoves. Fig.5 shows that theagreement between measurement and prediction isimproved when wall recesses are modelled. However,the agreement remains good when the room ismodelled as a simple rectangular volume. The effect ofwall recesses of depths less 0.5 m can be neglected.

859095

100105110115120125130135140145150

10 100 1000Frequency (Hz)

Soun

d pr

essu

re le

vel (

dB)

No chimney With chimney MeasurementsFig.5. Effect of wall recess on the sound field of a room

6. CONCLUDING REMARKS

Effects of construction, furniture and wall recess onroom response have been investigated by comparingfield measurement with FE models of variouscomplexities. It was found that the materialconstruction has a damping effect on the sound fieldwhich can be characterised by an equivalent surfaceabsorption coefficient. Below 100 Hz. the equivalentabsorption coefficient is 0.02 for plastered brick andconcrete, and 0.15 for lightweight cavity constructions.Furniture and wall recesses were found to have littleeffect at low frequencies and therefore need not beincluded in the FE model.

7. ACKNOWLEDGEMENTSThe authors gratefully acknowledge the financial support bythe Engineering and Physical Sciences Research Council ofthe United Kingdom.

8. REFERENCES1.J.R. Brooks and K. Attenborough, “The implication of measured

and estimated domestic source levels for insulation requirements”,Proceedings of I.O.A., vol 11, 19-27, 1989

2. C. Grimwood, “Complaints about poor sound insulation betweendwellings”, IOA Acoustics Bulletin, vol 20 (4), 11-16, 1995

3. S.P.S Maluski and B.M. Gibbs, “Application of a finite elementmodel to low frequency sound insulation in dwellings”, Journal ofAcoustical Society of America, vol 108 (4), 1741-1751, 2000

4 Melo G. ( 2001) ‘A Finite Element Model of Sound Absorption atLow Frequencies’, Euro-noise 2001

Sound Insulation of Lightweight Plasterboard Walls

J.Nurzyñski

Building Research Institute, ul. Ksawerów 21, 02-657 Warszawa, Poland

The paper presents an analysis of the results of laboratory measurements carried out on different samples of lightweight plasterboard walls. During the last two years over seventy partitions were tested on the same test facility under similar conditions. Different solutions of technical details were considered. The results show the influence on sound insulation of such factors as screw span, perimeter connection, rigidity of gypsum plates edges and joints, distance between studs or presence of additional elements joining studs. This factors can depend on workmanship, real building conditions or solution of specific wall structure details. It can be one of the possible reasons that the same partitions behave differently in different buildings.

INTRODUCTION Various factors influence the sound insulation of

lightweight plasterboard walls. Some of them are well known e.g. the thickness and number of plates on each side of the frame, sound absorbing material inserted into cavity or type and structure of frame. These elements are taken into consideration while designing partition of required sound insulation. However field measurements often reveal substantial differences between results obtained in the laboratory and in real building which are associated not only with flanking transmission. Even the laboratory tests conducted on semi identical samples of walls on different test facilities can reveal essential differences in sound insulation [1]. The analyse of collected results indicate possible reasons of such discrepancies.

MEASUREMENT CONDITIONS

In the last two years different kind of lightweight

plasterboard walls with different solutions of technical details were tested. All measurements were carried out on the same test facility with the same equipment and operator. Presented comparisons of results are for couples of walls erected by the same contractor with elements taken from one manufacturer and from one stock of material.

DISCUSSION OF RESULTS

Several partitions were tested with typical (20 cm)

span of screw joining panels to metal channels of frame and then with enlarged distance between screws (60 - 100 cm). The results obtained in case of typical single wall are compared in figure 1. The increase of distance between screws caused the increase of sound insulation in the range of medium and high frequency. The same effect was found in another eight samples of different single walls regardless of stud section, type of

absorbing material inserted into the cavity and number of plates on both sides of the frame. There are probably two main reasons causing such behaviour; different ratio of energy transmitted via studs connecting both leafs of plaster in each case and different rigidity of edges of plasterboard plates.

FIGURE 1 The influence of screw span on the airborne sound insulation of single plasterboard wall

Other result of enlarged distance between screws

was observed in case of double walls. The increase of screw span resulted in reduction of sound insulation in low frequency area. In case of double walls metal studs are separated and do not connect the panels fixed on both sides of the frame. Hence the screw distance does not influence this path of transmission (via studs). Reduced screw span restricted modal behaviour across the surface of panels in low frequency bands and resulted in increasing sound insulation in this range.

The influence of rigidity of gypsum plate edges themselves on sound insulation of lightweight wall can be observed when comparing results of measurements carried out immediately after plastering the edges of plates screwed to the frame (fresh soft gypsum plaster) and after a period of curing time when the connecting mortar is rigid and firm. It forms a sort of continuous rib on the perimeter of gypsum plate. Figure 2 shows the comparison of two curves; first obtained just after

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plastering joints and the second sixty minutes later. The drop in sound insulation after the period of curing time in the range of high frequency is conspicuous. Such tendency was observed in the case of another fourteen samples regardless of their frame type or structure. Similar effect was found when testing a single leaf of gypsum plate with completely free edges and after forming a small rib around the perimeter.

FIGURE 2 Change of sound insulation of lightweight plasterboard wall after period of curing time.

Another factor that can influence the sound

insulation of lightweight plasterboard wall is the connection of metal channels to the partitions surrounding considered wall. This path of transmission is of rising importance in case of double wall where both leaves of gypsum plates are connected to each other only on perimeter by surrounding structures.

FIGURE 3 Sound insulation of walls with different connection of metal channels to surrounding structure

Figure 3 shows the sound insulation curves of two

double walls which frames were constructed with 100 mm channels, 200 mm of mineral wool was inserted into cavity and two layers of 12.5 mm plasterboard were fixed on each side. In the case of partition “A” both parts of the frame are joined to concrete boundary of test facility opening on the same side of vibration brake (which is between totally

separated reverberant chambers of the laboratory). The frame of partition “B” is divided by the vibration brake. The difference in sound insulation is easily visible. The “lack” of critical frequency in case “B” is probably caused by the shortage of power in the source room in the high frequency range.

The distance between studs, which determine the dimensions of sub plates, also influence the acoustical

FIGURE 4 Sound insulation of lightweight walls with different distance between studs.

performances of the wall. Greater distance usually results in better sound insulation. An example is given in figure 4 where two characteristics of two single walls are compared. In both cases 25 mm thick gypsum plates were used and the distance between studs was respectively 31 and 100 cm. The second wall has significantly better sound insulation in medium and high frequency bands. A similar effect was noticed when comparing sound insulation of walls with 12.5 mm plates connected to each one (60 cm), and each second stud (120 cm). The difference in sound reduction index Rw was up to 5 dB.

CONCLUSIONS All presented technical details that influence the

sound insulation of lightweight plasterboard walls in the building can be solved differently. They depend on workmanship, design, building structure and specific solutions. It can be one of the possible reasons that the same wall behaves differently in different buildings.

REFERENCES 1. Pompoli R. “Inter-comparison of laboratory

measurements of airborne sound insulation of walls”, Final report, University of Ferrara 1997.

2. Pompoli R., Smith R.S. “Possible reasons for the discrepancy in the reproductibility results of the intercomparison of laboratory measurements of airborne sound insulation of walls”, presented at the 18th meeting of CEN/TC126/WG, Zurich, 1998.

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Evaluation of Room's Acoustical Qualities

B. Somek, S. Fajt, H. Domitrović

University of Zagreb, Department of Electroacoustics, Unska 3, HR-10000 Zagreb, Croatia In this paper subjective testing of room's acoustical qualities were investigated. Testing was made in the same conditions, with the same equipment, in two rooms with same dimensions, but with different acoustical properties. First room (seminar) was very reverberant, that is, not acoustically treated, and the other (studio) was acoustically treated. Many subjective parameters were investigated, and certain were chosen and modified. We derived definitions of the subjective acoustical parameters, and made scale of numerical marks. According to the testing results, room's acoustical qualities were evaluated and testing method was defined and suggested.

INTRODUCTION

In order to obtain the results of subjective examinations of the room’s acoustical qualities, it was necessary first to choose a suitable sound and speech sample. After detailed analysis we burned the test CD with chosen samples. The subjective tests were performed in two acoustically different spaces and 77 persons were included in both examinations. The examinees were students from Faculty of Electrical Engineering and Academy of Music in Zagreb. They had healthy hearing, the average age of 23 and fair musical education. The same sound system, consisting of a CD reproducer, an amplifier and a pair of loudspeakers, was used for reproduction of chosen sound samples in both rooms. There were no tone controls, but volume controls on the amplifier were on the same position during all tests. Before the listening was performed, every examinee got a short guide where all acoustical parameters considered in tests were described. It helps them to understand what to listen to during reproduction, rather than thinking about the meaning of the specific parameter. The judgment was accomplished by ranking each parameter with grades from 1 to 5. In fill-up questionnaire, given to each examinee, every rank for each subjective parameter was further described by a few words. There were several parameters that had two different descriptions for the same rank (e.g. “Reverberance”). Despite all preparations, there was still the possibility that some parameter was not clearly defined. If this was the case, the subjects were given the opportunity to leave such questions unanswered. The obtained results for each considered parameter are shown in Fig.1 to Fig.8. The numbers on y-axis depict how many examinees gave the corresponding grade (1 to 5) for specific subjective parameter. White bars are for seminar, while gray are for studio.

DEFINITION OF PARAMETERS

The space has acoustical intimacy if the reproduced sound left impression as it is really played in it. The suggested grades were: 0 (unanswered), 1 (far), 2, 3, 4, 5 (near). Reverberance has influence on “overlapping” of sounds. Reverberant places are usually described as “live”, while on the contrary the nonreverberant places are described as “dry”. The grades are 0, 1 (reverberant/dry), 2 (too small/ too large), 3 (very small/ very large), 4 (small/large), 5 (optimal).

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FIGURE 1. Results for intimacy.

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Timbre defines the perception of individual components of sound spectrum, at the listening position. It is influenced by sound system or room

acoustics. The grades are: 0, 1 (bad), 2 (weak), 3 (good), 4 (very good), 5 (excellent) Clarity expresses the separation of sounds from individual instruments among the other instruments playing at the same time. The grades are: 0, 1 (bad), 2 (weak), 3 (good), 4 (very good), 5 (excellent) Speech intelligibility means the intelligibility of speech or singing at the listening position. The grades are: 0, 1 (bad), 2 (weak), 3 (good), 4 (very good), 5 (excellent) Spectral uniformity parameter denotes balance of sound in the hearing band. This equilibrium depends upon loudness of instruments regarding the overall orchestra loudness. The grades are: 0, 1 (bad), 2 (weak), 3 (good), 4 (very good), 5 (excellent) Sound stage imaging describes how successful is the original sound picture from the stage preserved during reproduction. The grades are 0, 1 (unclearly), 2 (very muddy), 3 (muddy), 4 (slightly muddy), 5 (realistic). Ambience reproduction is the measure of ability for one acoustical space to simulate the ambience of the source space. The grades are: 0, 1 ((bad), 2 (weak), 3 (good), 4 (very good), 5 (excellent).

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FIGURE 4. Results for clarity.

CONCLUSION Upon given results of the subjective tests, it is clearly seen that there is no significant data spreading. Relatively small number of examinees left some

questions unanswered. This confirms that the questionnaire was well defined and described before listening tests, so it was possible to judge single parameter by ranking it with grade between 1 and 5. Here described evaluations are part of comprehensive subjective tests, performed in order to find the correlation between objective and subjective acoustical parameters of rooms.

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FIGURE 6. Results for spectral uniformity.

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Impact Sound Improvement on Lightweight Floors A Simplified Method

M. S. Kartousa, H. G. Jonassona

aAcoustic section of Acoustics, SP Swedish National Testing and Research Institute, SE-50115 Borås

It is well known that the impact sound improvement for floor coverings on lightweight floors is different from that on heavy concrete slabs. At present there is no test method which addresses this problem although an ISO working group has recently proposed a method using two different standardized lightweight floors. One problem with this proposal is that it is not cost efficient. Most laboratories have only one test opening and it is time consuming and thereby expensive to change test floors. A simpler test procedure that could be done without floor exchange would be preferable. The procedure described here involves a small lightweight construction placed on top of the standard concrete floor and could be easily set-up by two persons in about half an hour. The impact sound measurements are carried out in normal fashion. Tests that have been carried out on this test floor are compared with laboratory measurements on a lightweight joist floor. The results of this simple procedure are promising but not entirely conclusive and further work is needed before final evaluation.

INTRODUCTION

Several efforts have been made to introduce new test methods to make it possible to extend impact sound improvement according to ISO 140-8 [1] to be valid also for lightweight floors. 20 years ago Kaj Bodlund [2], carried out a Nordtest project ending in a method to use a standardized lightweight floor [3]. Although SP constructed such a floor it has never been used commercially. The main reason for this failure is that clients do not think that it is worth the effort considering the fact that floor coverings have only a minor effect when mounted on a lightweight floor. Just the floor exchange forth and back, including acoustic sealing, takes around 14 man-hours and also heavily affects the activity in the laboratory. (This can be compared with less than one hour for the proposed method.) The other reason is the fact that lightweight floor construct-ions can be very different from one case to another.

The Nordtest effort is now repeated within ISO and ISO TC 43/SC 2/WG 22 has now proposed a measure-ment standard including two different lightweight floors [4]. The ISO approach has, of course, the same advantages/disadvantages as the Nordtest approach. This paper will discuss the possibility to have a simplified approach based on the following:

• cost of the test rig and total man-time is low; • area of the test specimen is small, e.g. approx.

1-5 m2 and therefore two man portable; • improvement is measured on a construction where

only the upper board(s) of the real lightweight floor is simulated;

• upper board is mounted on wooden studs or steel bolt extensions placed directly on a standard laboratory concrete floor;

• different lightweight floors can be simulated easily and cheap.

SIMPLIFIED TEST METHOD The test floor consisted of a 2,4 m by 1,8 m large

22 mm thick floor chip board mounted on adjustable steel bolts c/c 600 mm. See fig. 1. The chipboard was identical to the upper layer of the joist floor. The bolts have to be trimmed into the same level so the whole chipboard floor is contact with the concrete. The whole procedure is simple and straightforward.

FIGURE 1. Test setup. Chipboard with parquet and under layer 2 visible.

A piece of thin rubber was placed underneath each bolt so the very small differences in surface level are absorbed. This also eliminates unwanted movement and unscrewing of the bolts due to vibration during measurement.

The test floor was placed on the 150 mm thick concrete floor normally used for impact sound tests according to ISO 140-8. The test floor was not fastened to the concrete slab, only weighted down with four 20 kg weights that ensured a satisfactory coupling to the concrete slab. By measuring with and without the floor covering to be tested, the impact sound improve-ment was obtained.

Three plastic floor coverings with soft foam layer and one parquet floor with two different under layers were tested.

The floor was exchanged to a lightweight joist floor from the NT project and the five test coverings were tested again for comparison.

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RESULTS The agreement between the improvement on joist

floor and chipboard test floor is within 4 dB at 50 to 160 Hz range and within 2 dB at 200 to 1250 Hz range. Above 1600 Hz the differences for the parquet floor and with two different under layers are around 4 dB. See fig. 2.

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The three plastic floor coverings tend a to get a large difference above 1,6 kHz. See fig. 3. As can be seen in table 1, the difference, expressed as ∆Lw,n, is insigni-ficant between the complete joist floor and the simpli-fied chipboard construction.

A more extensive testing is therefore needed before the mechanisms of the plastic covering behavior can be understood.

Table 1. ∆Lw,n comparison

Concrete floor

Wooden joist floor

Large chipboard

Plastic covering M1 15 1 1

Plastic covering M2 17 2 2

Plastic covering M3 19 2 2

Parquet under layer 1 17 5 4

Parquet under layer 2 16 3 3

CONCLUSIONS Bearing in mind there is no typical lightweight

construction that could be standardized, the proposed simplified method could be an interesting alternative to standardizing complete lightweight floor constructions to use in the laboratory. These test floors are easy and relatively cheap to produce. The construction does not have to be a single chipboard, although this adds to the portability advantage. This does not rule out the possibility to place the test floor on the concrete, trim it in height and then add another layer of chipboard or floor plaster. A floating floor test could also to be made with a small cement slab (1 m2). This would weight approx. 80 kg and be therefore portable.

ACKNOWLEDGMENTS The authors would like to thank Katrin Kohler who

performed most of the measurements.

REFERENCES [1] EN ISO 140-8:97 Measurement of sound

insulation in buildings and of building elements–Laboratory measurements of the reduction of transmitted impact noise by floor coverings on a heavyweight standard floor.

[2] Nordtest method NT ACOU 034 Ed.2 . Floor coverings, concrete or timber joist floors: Rating of impact sound improvement. ISSN 0283-7145

[3] Bodlund K., Carlsson C.A. Revision av nordtestmetod NT ACOU 034. Bestämning av golvbeläggningars stegljudsförbättringstal på trä- och betongbjälklag. SP-RAPP 1987:37

[4] ISO 140-11 Working Draft, Measurement of sound insulation in buildings and of building elements. Laboratory measurement of the reduction of transmitted impact noise by floor coverings on a lightweight standard floor.

FIGURE 2. Comparison of parquet floor with two different

under layers on a large chipboard test rig and joist floor.

FIGURE 3. Comparison of three plastic coverings on a large

chipboard test rig and joist floor.

Walking noise and its characterization

E. Sarradj

Institut fur Akustik und Sprachkommunikation, Technische Universitat Dresden, 01062 Dresden

The increasing distribution of hard wooden floors increased also problems with annoyance by the noise emission from

walking into the same room a person is walking. This shall be called walking noise emission to distinguish it from the

impact sound insulation topic which is still equally important. In the last years efforts were made to characterize this

noise and to quantitatively measure it. It turned out that at the moment no technical device is available which is able

to satisfactory simulate the walking noise. Therefore, an alternative method for walking noise measurement worked out

is topic of the presentation.

INTRODUCTION

Laminate and other types of hard wooden floorsbecame very popular in the last decade with a marketshare of over 10% in Western Europe. Not so popularis the annoyance by the noise emission from walkingon these floors. So the manufacturers made consider-able efforts to design low noise floors. The success ofsuch design can only be estimated by an appropriatetest method. As there is no agreed method avail-able this was the starting point for the developmentof such a method. Throughout this process the newvocable ’walking noise’ was introduced to distinguishthe problem from the related impact sound insula-tion. Also some theoretical insight into the genera-tion process of walking noise was gained.

THEORY

Laminate floor is usually laid on some elastic un-derlay. Therefore from a mechanical point of view itcan be seen as a plate on bedding. Cremer and Heckl[1] gave some expressions on bending wave propaga-tion on such plates. Damping in the laminate platemay be considerable high. So, in order to study walk-ing noise it is necessary to study the noise generatedby an impact on a damped plate on bedding. Theimpact of a body will generate two kinds of sound.First there will be a direct airborne sound generationthrough the sudden stop of the fluid mass (densityρ0) surrounding the impacting body. It is very diffi-cult to get an estimate of this sound as the impactspeed U0 must be known. Sound energy for a hardsphere of volume V sudden stopping is [1]

E =14ρ0V U2

0 . (1)

For other shapes of the body the equation must beslightly modified. The second kind of sound gen-

Table 1. Reverberation room measurements: SPL de-crease in dB compared to reference for 7 samples

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erated trough the impact is structure-borne sound,mainly bending waves in the plate. This sound isradiated back into the room and contributes to theoverall audible walking noise. The generation ofstructure borne sound depends on the elastic prop-erties of the laminate versus the impacting body, thearea mass of the laminate and the characteristics ofthe underlayer. The propagation of bending wavesin the laminate plate depends also on the damping.

TEST METHODS

There are two general possibilities to generatewalking noise - by a person walking or by a tech-nical device producing some impact sound. The re-producibility of the results gained by a specific testmethod is very important. Thats why two techni-cal devices - tapping machine and falling steel ball -were preferred at the beginning of the developmentof walking noise test methods.

For the test of the tapping machine some sampleswere laid in the reverberation room and the tappingmachine was used in the same way as for impactsound insulation measurements. A mean sound pres-sure level (SPL) was measured in the room using aspectrum analyzer. To test for meaningful resultsthese experiments were compared to the SPL from

a walking person on the same samples. Special carewas taken to make the results from the person notsubjective. With both methods the measured SPLswhere compared to a reference sample which was theloudest. The congruence of both methods was un-satisfactory, see table 1, and could not be improvedby calculatory alignment of spectra (fig. 1).

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A second method was tested using a steel ball(ø16 mm) falling from a defined height. The SPL wasaveraged over the first 100 ms after impact. Compar-ison to results from a walking person was bad again.

The reason why the technical devices failed istwofold. First, impact speed and mass are too dif-ferent from that of a walking person. The fact thatspectra could not be aligned indicates that nonlin-ear processes are of importance during the impact.Second, the direct airborne sound power is consider-able compared to the overall sound power, but it isdifferent for different objects1.

As a consequence from the results with technicaldevices a test method was set up which uses a walk-ing person to generate walking noise on a test facil-ity. Within a reverberant test room a special 60 mmconcrete basis with an even and an uneven surfacepart was laid. On top of this basis samples couldbe easily laid and tested. Testing procedure is asfollows: The walking person starts on a soft carpetand then walks onto the test sample. Measurementequipment is triggered on the first step on the testsample to average over 100 ms. For the overall resultthis is repeated at least 15 times. It was experimen-tally verified that this procedure assures a good re-producibility so that results from different measure-ment campaigns lie within an ±1 dB interval. The

1 rough estimates after (1) sound power level for 100 ms: 95 dBfor tapping machine, 75 dB heels shoe single step

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results for each sample are compared to a referencesample2 to get a measure of the enhancement. Fig.2 shows typical enhancements for different low noisetechnologies for laminates. Except from the lami-nates with thermoplastic damping treatment on theunderside the enhancement is poor or average.

SUMMARY

Currently there is no technical device availablewhich may generate walking noise in a manner com-parable to a person walking. A preliminary test fa-cility was set up to be able to estimate the perfor-mance of low noise laminate floorings. A walkingperson generates the noise which is then measuredtaking only one step. With this procedure a goodreproducibility can be achieved.

In the future a ”walking machine” should be con-structed which is able to reproduce meaningful re-sults also on other test facilities.

REFERENCES

1. L. Cremer, M. Heckl, Korperschall 2nd edition,Springer, Berlin, 1996.

2 reference is defined as 8.1 mm DPL laminate laid on PEfoam of brand Noppaschaum

Ratings adapted to subjective evaluation for impact andairborne sound and its application in building regulations–

a litterature surveyK. Hagberg

Division of Engineering Acoustics, Lund University, P.O. Box 118, SE-221 00 Lund, Sweden

Some acoustical requirements in standards or in Building regulations in several countries still suffer from seriousshortcomings. One of those is the sound insulation rating methods which could be better adapted to the subjective evaluation.During the last few years some important improvements have been made in the Nordic countries in national classificationstandards and building regulations.

During the development of the standards and the building regulations referred to above different aspects concerning measuresand levels were described and some former investigations were studied. Nevertheless, since the time available was restrictedit was difficult to collect a comprehensive view from all surveys made within the topic ”subjective ratings”. Current surveywill view aspects from former investigations and elucidate some mutual dependencies between these investigations. Themajor aim is thereafter to propose measures and requirement levels primarily concerning impact sound, appropriate to use inbuilding regulations and in standards at present. Finally, important topics for future research work will be discussed.

INTRODUCTION

Several investigations concerning the relationshipbetween objective measures and subjectiveevaluation have been carried out during the lastdecades and analysed further during the last fewyears [1, 2, 3, 4, 5, 6]. Current paper presents someimportant aspects on one of these investigations [4,5] and discusses some principles regarding theevaluation of the data in this investigation.

Investigation

From the impact sound level curves presented in thereports [4, 5] we have estimated the 1/3 octavelevels and recalculated all the data (original datawere not available). To ensure that the calculateddata are as correct as possible we compared allsingle numbers (L´n,w and the measure proposed byBodlund, Is) with those reported in the originalinvestigation. Furthermore, the correlationcoefficients were calculated and compared with theoriginal equations. The relationships became:

⟨L´n,w⟩ = 80,4 - 5,44 S [r = 75%, n = 22] (1)⟨IS⟩ = 85,9 - 5,43 S [r = 87%, n = 22] (2)

and the original relations

⟨L´n,w⟩ = 80,6 - 5,48 S [r = 75%, n = 22] (3)⟨IS⟩ = 86,3 –5,53 S [r = 87%, n = 22] (4)

where S is the subjective grading (1-7) and n is thenumber of data points. The small differencesbetween the equations might be explained bydifferences when rounding off, but also by the

difficulty to estimate correct 1/3-octave values foreach curve, since the curves in the reports sufferfrom precision. However, eqs (1) - (4) indicate thatwe are close to the original data.

From these “new” impact sound data some re-calculations and analyses are made. In researchworks presented by BRE [1,2] there are stated thatit is the floor structure above an apartment thatprimarily contribute to impact sound disturbance. Inthe research work [5] both horizontal and verticaltransmission are included. These facts give causefor concern particularly since the impact sound dataoriginating from horizontal measurements belongsto one “group” with low impact indexes while thevertical measurements belongs to a “group” withhigher indexes, see table 1 below.

Table 1. Measurement direction in relation to the valueof acoustic parameterMeasurementdirection [4,5]

Parameter Number ofdata points

Range ofacousticparameter

Horizontal L´n,w 9 37-49Vertical L´n,w 13 51-70Horizontal Is 9 43-56Vertical Is 13 59-72

Hence, all data from the horizontal measurementgroup exhibit mean values far below the minimumrequirements in current Swedish building code(L´n,w and L´n,w + CI,50-2500 ≤ 58 dB) and far belowthe values giving satisfactory (S ≥ 4,4 ⇒ Is ≤ 62dB) impact sound levels using Is [5]. Having theresearch work [1,2] in mind there is reason to saythat the results from the interviews might have beenaffected more by airborne sound for the “horizontalgroup” than for the “vertical group”.

Further calculations

Now, what happens if the two groups in Table 1 areseparated in the analysis? Only horizontalmeasurement gives the following relations:

⟨L´n,w⟩ = 59,8 – 2,59 S [r = 44%, n = 9] (5)⟨IS⟩ = 70,3 – 3,23 S [r = 53%, n = 9] (6)

which indicates weak correlation, both for the ISO717 single number and for the figure suggested byBodlund. Analysing only vertical transmission inthe original data:

⟨L´n,w⟩ = 70,1 – 2,48 S [r = 41%, n = 13] (7)⟨IS⟩ = 78,8 - 3,41 S [r = 81%, n = 13] (8)

and now, including two single measurementsoriginating from wooden structures with soft floorcoverings (section 4 and 6 in [4]) and in additionincluding data from two Swedish wood structureprojects [7] the following relations will appear:

⟨L´n,w⟩ = 67,4 – 2,22 S [r = 33%, n = 15] (9)⟨IS⟩ = 78,2 - 3,41 S [r = 77%, n = 15] (10)

The conclutions using these relations are that theL´n,w single number rating is even worse to use as ameasure describing the impact sound than what hasbeen shown earlier. However there is still one floorstructure in the supplementary study [5] that,according to my opinion, should be excluded sinceit gives strange and extreme subjective response. Itis a concrete floor structure covered with hardlinoleum floor covering in a sleeping room. Thehigh subjective ranking of this structure might beexplained by the fact that the impact sounds areemanating from a sleeping room. There are reasonsto suspect that, at least in this particular case,impact sounds from this room appear to a smallscale. Now the relationships will be

⟨L´n,w⟩ = 73,4 – 3,88 S [r = 60%, n = 14] (11)⟨IS⟩ = 78,9 – 3,60 S [r = 77%, n = 14] (12)

Naturally, excluding a floor structure with highsubjective grading and low IS only has minor effectwhen evaluating the two correlation relations (10,12). Analysing the relationship between the meanvalue of the ISO-measure introduced in severalnordic countries and the subjective response for thevertical case gives the following result:

⟨L´n,w+CI,50-2500⟩ = 73,4–3,80 S [r=77%, n =14] (13)

This relation (13) is plotted in Figure 1. The lowercurve represents the corresponding curve for thehorizontal case, eq. (14):

⟨L´n,w+CI,50-2500⟩ = 62,3–2,91 S [r=50%, n =9] (14)

FIGURE 1. ⟨L´n,w+CI,50-2500⟩ plotted against subjectiveresponse; × horizontal direction; ! vertical direction

SUMMARY

Correlation relations between objective measuresand subjective grading in the report [5] includedhorizontal and vertical impact transmission. In thispaper it is shown that by treating the verticaltransmission path separate, which would be moreattractive, the correlation coefficient becomesweaker, both for Bodlunds measure, IS, butparticularly for L´n,w. However, the new ISOmeasure including 1/3–octaves from 50 Hz isexactly as good as IS (to be compared with theresults in [6]) and, as far as we know today, thevalue should not exceed L´n,w+CI,50-2500≤56 dB(S=4,4). In an attempt to further improve the singlenumber evaluation we will include additional floorstructures to the study and with the “floor structure”approach try to find a measure even better adaptedto subjective response. In this deeper study we willalso try to illustrate some airborne sound aspects.

REFERENCES

1. C. J. Grimwood & N. J. Tinsdeall, Occupantopinion of sound ins. in converted and re-furbished dwellings in Engl. and the implica-tions for nat. Building Reg., Internoise (1998)

2. G.J. Raw & N.A. Oseland, Subj. Response toNoise Through Party Floors in ConversionFlats, Applied Acoustics 32 215-231 (1991)

3. J H Rindel, The rel. between sound insul. andacou. Quality in dwellings Internoise (1998)

4. K. Bodlund and L. Eslon, En kartläggning avljudklimatet i några moderna svenska bostäderTeknisk rapport SP-RAPP 1883:37 (1983)

5. K. Bodlund, Rating of sound insul. betweendwellings Techn. report 1985:01 (1985)

6. J. H. Rindel & B. Rasmussen, Assessement ofairborne and impact noise from neighbours,Internoise (1997)

7. P. Hammer, Reports TVBA-3000-3100 (1996)

L´n

2 3 4 4.4 5 6 735

40

45

50

55

60

65

70

75

56,7

49,5

L!"#$%$C&'()*+())$,-./ L!"#%C&'()*+())$01$0$23"4567"$72$1389:456;:$<:1=7"1:

New Partition Panel Construction Arrangements for betterSound Transmission Loss

M. Singha, b, K. K. Pujaraa, and V. Mohananb

aDepartment of Mechanical Engineering, Indian Institute of Technology, New Delhi-110 016, IndiabAcoustics Section, National Physical Laboratory, New Delhi-110 012, India

E-mail address: mahavir@csnpl.ren.nic.in (Mahavir Singh)

The acoustical characteristics of partition panel construction arrangements in theory are capable of providing extremely better valuesof sound transmission loss, but these better values are not usually obtained in practice, due to the presence of necessary soundbridges which often form a direct mechanical connection between the two panels that constitute the partition panel. A new partitionpanel construction arrangements has been developed to obtain for better sound transmission loss. This has been done by minimisingsound bridges (with practical examples of the bridging effect) and providing adequate cavity absorption. It has been possible toobtain transmission loss to provide an STC rating of 58, highly adequate for an inter dwelling partition panel, which is approximately10 rating points higher than the standard staggered stud wall panel, with little increase in total mass.

INTRODUCTION

The derivation of an expression for the soundtransmission loss of an ideal double panel is well-known[1, 2]. It can be shown [3] that the values of soundtransmission loss for diffuse sound field excitation are asshown in Fig. 1. It is assumed here that there are nomechanical connections between the two panels whichconstitute the structure, and that the two panels obey themass law over the frequency range of interest. Fig. 1shows clearly three distinct frequency regions in whichthe sound transmission loss increases at rates of 6 dB, 18dB and 12 dB per octave, respectively. These rates ofincrease will be modified if the effects of coincidencehave to be included in the problem. The differencebetween these predicted values and those predictedaccording to the mass law indicate the possibility of greatincreases in sound transmission loss for double over andabove single panel structures.

FIGURE 1. Sound transmission loss of a general double panel.

RESULTS AND DISCUSSION

Experimental evidence to support this hypothesis hasbeen obtained by measuring the sound transmission loss

of a double panel in which the individual panels werecompletely isolated. The experiments were conducted inthe suite of reverberation chambers at the AcousticsSection, National Physical Laboratory of New Delhi inaccordance with the requirements of ASTM E90-1990,and of ISO 140/III 1978(E) and STC with ASTM E413.The double panel was placed in the opening (640 x 940mm). The edges of the cavity were sealed. To eliminatethe effects of coincidence from the study, the panels usedwere 3 mm and 6 mm hardboard. Results of theexperiments are shown in Fig. 2.

In the absence of absorption, curve (a) of Fig. 2 showsthat the strong acoustic coupling between the panelsresults in a single panel performance at frequencies lessthan the first cavity resonance perpendicular to the planethe panels. At higher frequencies, the phase of the soundpressure varies over the thickness of the cavity and theacoustic coupling is weaker. In that frequency range,sound transmission loss is seen to increase and behavemore like a double panel, although the predicted valuesare not attained. The introduction of a 50 mm layer ofglass fibre batts (48 kg/m3) across the entire cavity width

FIGURE 2. Effect of cavity absorption on the soundtransmission loss of a double panel.

(curve (b)) produces a remarkable improvement in soundtransmission loss. The agreement between theory andexperiment in this case is good.

The method used to determine the reduction in soundtransmission loss of a double panel due to the insertion ofa number of sound bridges is to sum the acoustic powerradiated by the action of the bridges and that radiated bythe ideal, isolated panel. This is performed [3] with theresult as shown in Fig. 3. The increase in soundtransmission loss above the calculated mass low for thecomplete structure (�TL in the Fig. 3) depends on whetherthe bridges are point or line connections. It also dependson the number of such bridges and the critical frequencyof the panels. Better values are obtained with few bridgesand flexible panels.

FIGURE 3. Effect of sound bridges on the sound transmissionloss of a general double panel.

A practical example of the bridging effect is shown inFig. 4 for the case of a single wood stud wall (studs 600mm o.c.) with 15 mm and 9.5 mm gypsum board panels.The improvement of point over line connections is on theorder of 5 to 8 dB in the mid-frequency region. It is to benoted that there is a good agreement between theory andexperiment.

FIGURE 4. Sound transmission loss of double panel withsound bridges.

The presence of sound bridges, together with inadequatecavity absorption, are the main reasons for the poor resultsoften obtained from double panels. Particular care musttherefore be taken with these two factors in thearrangement of better sound transmission loss structures.An example is shown in Fig. 5, which is basically a

staggered wood stud partition with gypsum board.Laminated gypsum board is used to increase the panelmass without increasing the stiffness. Both panels aremounted on points spaced 610 mm apart, althoughevidence now exists to suggest that point connections arerequired only on one side. Additional benefits areobtained by mounting the laminated panel on 6 mm PVCpads that act as vibration isolators. It is expected that padsmade of polystyrene will be equally effective. It ispossible to nail through the isolators without too great adetraction in acoustic performance. The panelconstruction arrangement geometry shown in Fig. 5provides an STC rating of 58, which is ideal for an inter-dwelling partition panel. This is approximately 10 ratingpoints better than the standard staggered stud wall, withlittle increase in total mass.

FIGURE 5. (a) Modified staggered stud constructionarrangement; (b) Sound transmission loss of modified staggeredstud construction arrangement.

The same construction arrangements can be applied tomany other structural types, both new and conventional, toprovide STC ratings of 70 to 80. The associated costs andarrangements are comparable to those of present dayconstruction arrangements which provide greatly inferioracoustical performance.

REFERENCES

1. Singh, M., Pujara, K.K., and Mohanan, V., “A physical approach ofsound transmission through panels,” Accepted in Indian J. PureAppl. Phys.

2. Fahy F.J., in Fundamental of noise and vibration, edited by F.J.Fahy and J.G. Walker, E & FN Spon, London, 1998, pp. 293.

3. Work presently being performed under supported by India Gypsumlimited. To be published. M. Singh, et al.

Problems of measurement and evaluation of noise from building equipment according to new European standard

pr EN ISO 16032

M. Mirowska M. Niemas

Department of Acoustic, Building Research Institute, 02-656 Warsaw, 21 Ksawerów St.

In paper the noise measurement method proposed in standard pr EN ISO 16032 is discussed. On the basis of test measurements of equipment noise penetrating into rooms the problems of evaluation of noise levels are presented. The results of noise measurement in dwellings carried out according to pr EN ISO 16032 and pr EN ISO 10052 are compared. The proposal of assessment method of octave noise spectrum is given.

INTRODUCTION

At present in the work group CEN/TC 126/WG 1 projects of two new European norms of measurement of noise fro m service equipment in building are discussed: pr EN ISO 10052 [1] and pr EN ISO 16032 [2] Survey method according to pr EN ISO 10052 [1] consists in measurement of sound level A for two positions of microphone, including one a corner, for 3 operating cycles of an equipment. The measurements are carried out by sound level meter, and result of measurement is an average from 3 readings sound level A or C (equivalent and/or maximum) standardized or normalized, without correction for background noise.

In special cases, one should use engineering me-thod according to prEN ISO 16032 [2]. This method consists in measurement in 3 positions of equivalent and/or maximum SPL in octave bands in the range from 31.5 Hz to 4000 Hz for analogous operating cycles as in the survey method. The sound level A and C is calculated from averaged and corrected results of measurement (correction of background and room absorption). Is yet the engineering method, recom-mended by pr EN ISO 16032, really more exact? Answer to this question has been searched on the way of test measurements according to this methods.

RESULTS OF TEST MEASUREMENTS

On the grounds of recommendations of projects of the standards pr EN ISO 16032 and pr EN ISO 10052 attempts were undertaken to perform measurements in some inhabited buildings. The measurements were effected with an application of mobile frequency analyser SVAN 912 or B&K2231. For recommended operating cycles a level of sound A and C – of noise and acoustic background were registered in selected measurement points (by means of "plot" function). In order to assign the maximum and equivalent SPL in

octave bands multi-spectrum was recorded for whole operating cycle.

Analysis of measurement results permits to state, what follows: 1. At low levels of nois e (20-30 dBA), that mostly appear in habitable rooms, the measurement results are in great measure affected by instrument own noise, and in particular, accidental levels registered at switch-on-moment of the instrument (fig.1). It is necessary to average levels from restricted section of time in order to define a level that really appears in the given point. After use of such an averaged time restricting procedure much lower levels are attained which determines final estimation of noise (see fig. 2).

0

5

10

15

20

25

30

35

0 10 20 30 40 50 60 70 80 90 100 110Record number

SP

L, dB

1000 Hz

FIGURE 1 The multispectrum SPL registered during mea-surements of acoustic background in a room

2. Whereas for continuous noises of settled levels it was possible to register spectrums for some work sections of 30 seconds each, so in case of transient noise it was never possible to perform measurements for more than 3 operating cycles. It was too much time -consuming and arduous both for inhabitants of the building, in which noise was measured as for measu-ring team. For measurements of noise of levels up to 30 dBA, it is practically very difficult for longer sec-tions of time, in duration of operation cycle of tested equipment, to assure circumstances without accidental acoustic disturbances. And the operating cycles are in many cases comparatively long e.g.: driving with a lift

up and down with stops and opening the door on every floor, or filling a bathtub to half of its capacity.

0

5

1 0

1 5

2 0

2 5

3 0

3 5

4 0

3 1 . 5 6 3 1 2 5 2 5 0 5 0 0 1 0 0 0 2 0 0 0 4 0 0 0 8 0 0 0 L A

f ( H z )

SPL,

dB

A 7

A 1 5

FIGURE 2. Equivalent SPL for background noise in a room, averaged for the whole section of observation time (A7) and for restricted section (A15).

3. Comparing results of measurements of noise level, executed simultaneously for the same equipment according to pr EN ISO 10052 and according to pr EN ISO 16032 (fig. 3) one can notice, that in most cases, sound levels A defined according to prEN ISO 16032 are lower than those defined according to prEN ISO 10052. The differences may result as well from regard to influence of acoustic background, as from another way of regard to correction A, i.e. fluent correction comprising the whole band of acoustic frequency, so as it is carried out in case of direct measurement or only for the middle frequencies of the octave bands, so as it takes place in case of the methodology proposed in prEN ISO 16032.

29,3

26,1

36,8

33,8

27,6

32

27

35 35

28

2 0

2 5

3 0

3 5

4 0

4 5

f a n 1 p u m p f a n 2 a i r - c o n t r a n s f

L A (d

B)

c a l m a e s u r

FIGURE 3. Comparison of sound level A values from mea-surements according to prEN ISO 10052 (1) and according to prEN ISO 16032 (2)

4. Having analysed measurement results received in the engineering method for more than ten sources a statement has been made that in all cases a notice should be recorded in report, that the result is influenced by a background noise. Even in cases, when difference between noise and background amounted to 10 dB A, for spectrum of noise always frequencies happened for which the difference between SPL of noise and background was smaller than 4 dB.

5. Controversial is also acceptance of an average energetistic in time and space spectrum of background

and noise. If measurements are made for hygienic purposes (exposure to noise), a spectrum of noise from a measure point of the highest levels should be taken, and not an average one in a room.

4. Another deficiency of the pr EN ISO 16032 consists in not clear definition of the way of assessment of maximum level spectrum – are those the maximum values for each frequency from the whole cycle or is it a spectrum corresponding with maximum level A or C. Neither is it clearly said, which partial can be acknowledged as tonal.

ASSESEMENT OF PROPOSAL METHODS

Verification measurements of noise penetrating to rooms from equipment installed in buildings showed, that measurements according to engineering method are not exact and they do not ensure an estimation of noise, more reliable than measurements according to EN ISO 10052. Instead, they are more difficult, more labour-consuming and they require use of analysers in real time of low levels of own noise with large "buffer" of memory. It would not be a bigger difficulty, if with use of the same analysers measurements were executed for tierce band in wide range of frequency, having regarded even the range of low frequency noise from 8 Hz measurement frequency, and results would surely be more exact.

As sufficient for control purposes could be considered result of measurement of noise spectrum from one loudest position, situated in places of regular stay of inhabitants and background spectrum be registered from section of time without accidental noises (not average ones). For comparison with noise limit, levels A and C counted from full measure range (e.g. 8-10000 Hz) and eventually normalized or standardized according to EN ISO 10052.

In controversial cases an evaluation of noise could be made by comparison of measured spectrum of noise and spectrum of background (without any correction). Our experiments prove that if in spectrum of noise partials appear, for which SPL is by 5 dB larger than the level of background - thus a noise can be audible and arduous. To estimate noise spectrum the characteristic LA10=10 –kA can also be used - for spectrum in tierce bands or LA15=15 –kA – for octave bands - corresponding with the level LA ≈25 dB thus corresponding with relatively comfortable acoustic conditions

REFERENCES

1. Pr EN ISO 10052. Acoustics – Field measurement of airborne and impact sound insulation and equipment sound – Survey method

2. Pr EN ISO 16032. Acoustics - Measurement of Sound Pressure Level from Service Equipment in Building – Engineering method.

Numerical Study on Sound Insulation Performance ofMembranes with Additional Weights

T. Oshimaa and T. Sakumab

aDepartment of Civil Engineering and Architecture, Niigata University, 8050 Igarashi-Ninocho, Niigata City, JapanbInstitute of Environment Studies, the University of Tokyo, 7-3-1 Hongo, Bunkyo-ku, Tokyo, Japan

Sound insulation performance of a MAW (membrane with additional weight) is tuned through arrangement of weights. A study isconducted to compare sound transmission through MAW with PWs and LWs (point- and line-shaped weights). The results indicate1) the LWs without stiffness show broader-band transmission loss and higher peak frequency than the PWs, 2) the width of finiteelements affects the frequency step of small dips in transmission loss for LWs, 3) longitudinal directivity of incident angle is notobserved as to the PWs, whereas for the LWs the maximum transmission loss is obtained at the angles along the LWs.

1 INTRODUCTION

It is known that the MAWs (membrane with addi-tional weights), developed by Hashimoto et al.[1], havehigh sound insulation performance based on a mech-anism different from the mass law. The performanceof a MAW is flexibly tuned to meet the needs of usersthrough optimizing weight arrangements including theshapes of the weights. The authors conducted a numer-ical study to make comparison between the insulationperformance of MAWs with LWs (line-shaped weights)and PWs (point-shaped weights) using FEM- and BEM-based wave analysis method.

2 NUMERICAL METHODS

Consider a MAW of 0.9m×0.9m mounted on an in-finite rigid baffle as illustrated in Fig. 1. The vibro-acoustically coupled matrix equation is formulated byapplying a structural FE (finite-element) model to the vi-bration system of MAW and a boundary element modelto sound fields[2]. The elements of inertance matrix[M]for MAWs are written as follows

Mi j = δi j

(mi +

13

N

∑k=1

ρikl ik

)+

16

ρi j l i j +Mmi j(1)

wheremi is the mass of the PW on thei-th FE node,ρi jand l i j are the linear density and the length of the LWbetween the neighboringi-th and j-th nodes,Mmi j

is the

(i, j) element of the membrane inertance matrix. Thefield incidence transmission lossT Lf is calculated fromthe weighted average of 120 incident directions[3].

3 RESULTS AND DISCUSSION

Tab. 1 shows the properties of MAWs. The MAWsare uniformly divided into 30× 30 FEs (the width of aFE∆ = 0.03m). The PWs and the LWs are placed on theFE nodes and the grid lines respectively.

infinite rigid baffle

planarincident wave

0.9m

0.9mMAW y

x

z θ

ϕ

O

FIGURE 1. Geometric configuration of sound transmissionmodel for MAW.

Table 1. Properties of MAWs.

Point (PW) Line (LW)

m = 0.02 kgMass / Density

Arrangement

ρ = 0.214 kg/m

Common properties: weight spacings d = 0.09 m,aereal density σ = 1.0 kg/m2, tension T = 4000 N/m

x

y

Shape of weights

Characteristics ofTLf for the PWs and the LWs. Fig.2 comparesT Lf for the PWs, the LWs and the masslaw of the same areal density including the additionalweightsσ = 3.0kg/m2. Both types of weights showincreases inTLf in lower frequencies compared to themass law. The LWs show higher peak frequency, lowerpeak level and broader transmission loss than the PWs.

The dips seen at about 12Hz step for the LWs areconsidered to be because the definite width∆ of a dis-cretized LW makes the LW strained by tensionT∆. Thusthe LW has the natural frequency step

fn ≈ (1/2L)√

T∆/(ρ +σ∆) (2)

where L is the length of a LW. For the present casewe obtain fn = 12.3Hz, which corresponds to the fre-quency step. The vibration displacement under nor-

-20-10 0

10 20 30 40 50

0 100 200 300 400 500

Tra

nsm

issi

on lo

ss [d

B]

Frequency [Hz]

PWLW

Mass law

fd fp

FIGURE 2. Comparison ofTLf in PWs and LWs.

0

0.3

0.6

0.9

0 0.3 0.6 0.9x [m]

y [m

]

0 0.3 0.6 0.9x [m]

-60

-40

-20

0

20

40

60

Dis

plac

emen

t [µm

]

Real Imaginary

FIGURE 3. Vibration displacement of MAW with LWs atfd.

mal incidence condition at one of the dip frequenciesfd = 212.5Hz, shown in Fig. 3, illustrates a coherentvibration of the LWs. The mechanism is confirmed bythe Fig. 4, which compares the same condition exceptthe 60×60 FE division (∆ = 0.015m) of the MAW. Theobserved frequency step of about 8Hz conforms to theequation (2).

Fig. 5 shows the vibration displacement of LW atpeak frequencyfp = 244Hz under normal incidencecondition. The vibration modes of the LWs are similarlyobserved whereas membrane is almost at rest.Directivity of transmission loss against incidentwaves. Fig. 6 contrasts oblique incident transmissionloss at incident angles(θ ,ϕ), TLθ ,ϕ , for the PWs andthe LWs at each peak frequency. TheTLθ ,ϕ for the PWsis not affected byϕ (the longitudinal direction), whereasfor the LWs the maximum transmission loss is obtainedatϕ = 90◦ (the angles along the weights).

4 CONCLUSIONS

A numerical study on weight shapes for MAWs indi-cates that theT Lf of LWs show higher peak frequency,lower peak level and broader band than PWs and thatLWs have longitudinal directivity in transmission lossagainst the incident waves. The small dips seen in LWsremain to be resolved because the dips make frequency-averaged transmitted energy which affect octave-band

0

10

20

30

40

140 180 220 260 300

Tra

nsm

issi

on lo

ss [d

B]

Frequency [Hz]

30x30 FEs60x60 FEs

FIGURE 4. Comparison ofTLf in 30×30 and 60×60 FEs.

0

0.3

0.6

0.9

y [m

]

Real Imaginary

0 0.3 0.6 0.9x [m]

0 0.3 0.6 0.9x [m]

-1.2

Dis

plac

emen

t [µm

]

-1-0.8-0.6-0.4-0.200.20.40.6

FIGURE 5. Vibration displacement of MAW with LWs atf p.

38 40 42 44 46 48 50

0

30

60

90

0 26 52 78θ [deg.]

ϕ [d

eg.]

TLθ,ϕ [dB]

x

yϕ

incident wave

26 28 30 32 34 36 38 40

incident wave

x

yϕ

0 26 52 78θ [deg.]

TLθ,ϕ [dB]

FIGURE 6. Effect of incident angles(θ ,ϕ) to TLθ ,ϕ for thePWs (left) and the LWs (right) at each peak frequency ofTL f .

transmission loss dependent of the FE size.

REFERENCES1. Hashimoto, N., Katura, M., Yasuoka, M. and Hujii, H.,J.

of Archit. Plann. Environ. Engng, AIJ410, 1–8 (1990).2. Sakuma, T. and Oshima, T., “Prediction of sound insulation

of partitions by vibro-acoustic analysis”, inProc. AutumnMeet. Acoust. Soc. Jpn., edited by ASJ, Tokyo, 2000, pp.687–688.

3. Sakuma, T. and Oshima, T., “Application of a vibro-acoustic method to prediction of sound insulation perfor-mance of building elements”, inProc. of Internoise, Hague,2001 (in press).

Analysis of problems to express uncertaintiesof building acoustic measurements

H. Goydke

Physikalisch-Technische Bundesanstalt (PTB), Bundesallee 100, 38116 Braunschweig, Germany

The application of international harmonized measurement standards will be mandatory in the field of building acoustics in thenear future in the EU member countries in order to facilitate trade across the borders. The question of the uncertainty of mea-surement results characterizing the acoustical performance of building products is getting growing importance therefore. In ISO140-2 of 1991 only preliminary and insufficient precision data are given regarding the measurements results when the ISO 140measurement standards are applied. To start the revision in ISO/TC43/SC2/WG18 a comprehensive analysis of a lot of problemsis necessary. The expression of precision data by the terms "repeatability" and "reproducibility" will have to be transferred to thequite different procedure according to the "Guide to the expression of uncertainty in measurement" (GUM). Recent inter-labora-tory comparison measurements enable the analysis of these problems to be based on realistic figures.

INTRODUCTION

Most of the international building acoustic mea-surement standards are meanwhile EN ISO harmo-nized standards. In any case where a manufacturerattests that a product conforms to such a harmonizedEuropean standard it has to be accepted in everymember state of the EU that the product satisfies therequirements of the “Building Product Directive”.The question of comparability of building acousticmeasurement results therefore gets growing impor-tance. Testing laboratories will have to be accreditedaccording to EN ISO/IEC 17025: 1999 [ 1 ] whichdemands: “Testing laboratories shall have and shallapply procedures for estimating uncertainty of mea-surement.” The meanwhile widely internationally acceptedbasis for the evaluation and statement of the mea-surement uncertainty is the "Guide to the expressionof uncertainty in measurement" (GUM) [ 2 ]. ISO TC43 “Acoustics” in December 2000 has urgentlydemanded that GUM shall be applied in the field ofacoustic measurements as well. SC 2 “BuildingAcoustics” will have to analyze the present situationas a first step to which this paper will contribute someconsiderations.

Uncertainty statements of ISO 140-2

The harmonized ISO 140 comprises in the atpresent 11 parts measurements of airborne and im-pact sound insulation of building elements in thelaboratory and as well of sound insulation betweenrooms and of facades in the field. Measurementsresults are to be taken in 1/3-octave bands (partlyalternatively in octave bands) within 100 Hz to 3150Hz (mid-band frequencies) respectively in the en-larged range from 50 Hz to 5000 Hz.

As Part 2 in ISO 140 is included the “Determina-tion, verification and application of precision data”.But this standard of 1991 says explicitly that thevalues given are tentative and incomplete and it isespecially to be mentioned that the values given arerelated to the different parts of ISO 140 before therevision of that parts took place which had, besidesothers, the aim to improve the precision of the mea-surement methods. The procedure for the evaluationof precision data is based on ISO 5725 [ 3 ]. Accordingly the most important figure to charac-terize the precision of a measurement result is the“Reproducibility value R”. R is determined from theresults of inter-laboratory tests and is defined as the“value below which the absolute difference betweentwo test results obtained under reproducibility condi-tions may be expected to lie with a probability of 95%”. Such conditions are: The same method on identi-cal test material in different laboratories with diffe-rent operators using different equipment. The repro-ducibility value R is given by

228,28,2 LrR sssR +== (1)

sR is the reproducibility standard deviationsr

2 is the mean of the within-laboratory variances taken over all participating laboratoriessL

2 is the between-laboratory variance taken over all participating laboratories The factor 2,8 takes statistically into account that Ris applied to differences between two single results.The following interpretation of R is valid: “In a singlefrequency band, the difference between two singleresults on identical test material reported by twolaboratories should differ by more than the repro-ducibility value R on the average not more than oncein 20 cases” and “between two sets of measurementcomprising the values in all frequency bands thedifference between results throughout the whole

frequency range will on average exceed not morethan once the reproducibility values”. It is obvious that the use of reproducibility valuesmeans to characterize precision with only one singlevalue having in view worst case situations andlooking to the maximum variability in test results.But meanwhile it is questionable whether thisapproach with its opaque character due to a obviouslylarge number of influences is satisfying.

GUM in relation to ISO 140-2

GUM [ 2 ] emphasises that it is a purpose of theguidance “to promote full information on howuncertainty statements are arrived at”. From the spe-cific terms of the guide the principal of the procee-dings can be seen and the difference to ISO 140-2 isobvious. These terms are 1. “Standard uncertainty” asthe result of a measurement expressed as a standarddeviation u; 2. “Type A evaluation” by the statisticalanalysis of series of observations; 3. “Type B evalua-tion” by means other than statistical analysis; 4.“Combined standard uncertainty” equal to the posi-tive square root of a sum of the variances or co-vari-ances of a number of quantities weighted accordingto the variation with changes of these quantities; 5.“Expanded uncertainty” is derived by multiplicationof uc with (6.) the “Coverage factor” k which is typi-cally 2 (95 % probability). In [ 4 ] it was shown by the example of the contri-bution of a sound level meter to the uncertainty ofnoise measurements that the GUM-procedure is to beapplied in the acoustic field as follows: The modelfunction which is to be found is in this case:

∑=

+=n

iiAA LL

1

' δ (2)

AL = sound pressure level; 'AL = sound level meter

reading; iδ = correction due to imperfect per-formance of the sound level meter. The uncertainty uis attributed to LA as follows:

( ) ( ) ( )∑=

+=6

1

2'2

iiAA uLuLu δ (3)

In a in a “uncertainty budget” at first the contributionof u(L’A) is evaluated where the error limits of theverified instrument are taken as standard uncertainty(0,02 dB) which is multiplied with a sensitivity factorof 1 and the probability distribution (normal) is takeninto account. u (L’A) = 0,02 dB. In the same way 6contributions u (δi) are calculated: drift δ1 = 1,17 dB,

directivity δ2= 0,20 dB, frequency weighting δ3= 0,28dB, linearity δ4= 0,40 dB, differential linearity δ5=0,23 dB and rms detector corr. δ6= 0,29 dB. It fol-lows: u(LA)= 0,7 dB and U(LA)= 1,4 dB.

Obviously in the R-values of ISO 140-2 the stan-dard uncertainty values according to GUM are in-cluded as “Type A evaluation”. Assuming no otheruncertainty contributions the conversion to the GUM“expanded uncertainty" therefore is simply done byapplying the factor of 0,714. But most important:those converted values can be taken as a part of theuncertainty budget and easily supplemented byadditional components. In Table 1,A R-values of ISO140-2 according to 140-3 (laboratory measurementsof airborne sound insulation) are listed, Table 1,Bshows the values converted to GUM U(R). Table 1,Cgives an impression of new data according to therevised ISO 140-3: In 1999 10 laboratories inGermany have performed comparison measurementsof the same test objects in a PTB test-facility but withtheir own equipment and staff. Those new results willenable in future to state uncertainty even for singlenumber quantities consisting of two parts to whichdifferent frequency bands deliver contributions. The conclusion of this consideration is that ISO140-2 and GUM are obviously not so incomparableas it may seem. The GUM procedure is flexibleenough to enable the transformation of the tentativeISO 140-2 values and to combine them with data ofrecent investigations. Table 1. Reproducibility values R according to ISO 140-2 and GUM expanded uncertainty values U

1/3-octavebands[ Hz ]

A: R of 140-2acc.140-3 (old)

[dB]

B: U = 2 sR

acc. to GUM[dB]

C: U = 2 sR for140-3 (new)

[dB]50 2,663 4,880 2,2

100 9 6,4 2,8125 8,5 6,1 3,2160 4,3 3,6200 3,9 2,4250 5,5 3,9 1,8315 4,5 3,2 1,4400 4,5 3,2 1,2500 4 2,9 0,8630 6 2,5 0,8800 5,5 2,1 0,61000 2,5 1,8 0,81250 3 2,1 0,31600 3,5 2,5 0,52000 3,5 2,5 0,82500 3,5 2,5 1,03150 3,5 2,5 0,84000 0,85000 0,8

REFERENCES1. EN ISO/IEC 17025: 2000; General requirements of the

competence of testing and calibration laboratories2. GUM, 1993, corr. and reprinted 1995, ISO, Geneva3. ISO 5725: 1986; Precision of test methods4. Brinkmann, K., J. Acoust. Soc. Am. 108, p. 2550

labs

situslabsitu T

TRR

,

,lg10−=

How will heavy walls be measured in future in test facilitiesaccording to ISO 140

A. Schmitz*, H-M. Fischer**

*Institute of Building Physics, Mainstrasse 1, D-45478 Mülheim a.d. Ruhr, Germany** University of Applied Sciences, Department of Building Physics, Schellingstrasse 24,

D-70174 Stuttgart, Germany

In future a new concept is needed where the laboratory measuring results of sound insulation of building elements are in directagreement with the new European calculation methods (EN 12354). Especially concerning heavy building elements the totalloss factor (TLF) must be included in the laboratory measurement procedure. If the measured sound reduction index (SRI)results are referred to a reference TLF differences due to the different test facilities are eliminated. It is shown that the TLF infield does not vary in that way as theoretically expected. Therefore a mean value of the field TLF is proposed to be thereference one. Then the input data for the calculation model obtained by laboratory measurements are directly in closeconnection to field measurements. Applying the simplified model of EN 12354, which does not provide in-situ correction, withTLF corrected laboratory values as input data leads automatically to laboratory consistent and consistent field results.

INTRODUCTIONIt can be shown theoretically and experimentally, thatthe actual sound reduction index (SRI) of a buildingelement in general depends on its total loss factor(TLF). The European calculation model for airbornesound transmission (EN 12354-1) takes this effect intoaccount by using a so-called in-situ correction tocalculate the SRI in real buildings. For the direct soundtransmission of the building element the followingequation is given,

(1)

where Rsitu and Ts,situ are the SRI and structuralreverberation time respectively in the real building andRlab and Ts,lab are these values under the laboratoryconditions. This formula in general allowsmeasurement results obtained in the laboratory to betransformed to field conditions. The meaning of thistransformation procedure is highly important if heavyelements are concerned. The energy flow and,therefore, also the TLF of heavy elements is muchdetermined by its boundary conditions. Even in testfacilities which are in accordance with the current ISO140-1 standard the given boundary conditions are quitedifferent. Depending on the construction of the testfacility, the TLF of the measured building elementmay vary very much. As this dependency is known,the measurement of TLF as an additional informationis recommended in the current version of ISO 140-3concerning heavy elements, but is not yet a must..

NEED FOR A RESVISION OF THESTANDARD

The meaning and importance of the TLF was recentlyshown by experimental work. An Inter-laboratory Test(ILT) was carried out in Germany, where the SRI of aheavy wall was measured in 11 different test facilities[1]. The test object was a calcium silicate wall(m´=440kg/m2). In addition to the SRI, the total lossfactor was measured using a measuring methoddeveloped by PTB. The main results are shown in Fig.1 and can be summarized as follows:

52

54

56

58

60

62

A B C D E F G H I K L M

Wei

ghte

d So

und

Red

uctio

n R

w

Participants 52

54

56

58

60

62

A B C D E F G H I K L M

Wei

ghte

d So

und

Red

uctio

n R

w,c

Participants

FIGURE 1: ILT results , SRI of 11 laboratories a) not converted b) converted

The spread of the measured Rw values is(Rw=54...61 dB). This spread was reduced sig-nificantly by application of a conversion where theresults were referred to a reference TLF according toequation 2.

Using this reference value all results were referred toan equal energy flow situation. In comparison with the

slabf

refslab

refs

labslab T

TR

TT

RR ,

,

,* lg10lg10 −=+=)100/(lg3,34,12lg10 f−−=η (3)

current version of ISO 140-3 this procedure leads toanother understanding of the SRI measured in testfacilities, because the properties of the test facility candirectly be taken into account.

In addition a direct relationship can be seen betweenthe proposed measuring procedure and the CENcalculation model.

DETERMINATION OF THEREFERENCE VALUE

At first glance the choice of the reference value forapplication of the correction to laboratorymeasurements seems to be unimportant, but if themeasured values are to be used as input data for thecalculation model, this reference value will getimportant. The simplified model of EN 12354 does notprovide in-situ correction. The values obtained bylaboratory measurements are used directly. For goodagreement between laboratory and fieldmeasurements, the reference value of TLF should bechosen to be as close as possible to TLF values foundin the field. The question to be answered therefore are:“How large is the spread of TLF in the field?” and“Can field situations be represented by only a singlereference value”.

FIGURE 2: TLF measurements, spread and meanvalue of 210 building elements measured

In order to validate the CEN calculation models and toanswer this questions, a large number of measurementsand investigations were carried out at the University ofApplied Science in Stuttgart [2]. Measurements of theTLF of about 210 different building elements (wallsand floors, m´=63...486kg/m2) were carried out in thefield. The first result of the latter investigation was thatthe measured TLF was not in good agreement with thevalues calculated according to EN 12354-1 (Annex C).

Another interesting result is shown in Fig. 2. Thespread of the measured TLF is quite smaller thanexpected. The mean value can be given as a regressioncurve according to equation 3.

Fig. 3 shows the difference between TLF valuescalculated according to equation 3 and measured TLF.The mean value of the difference is 0,29 dB. Even thestandard deviation (2,12 dB) seems to be tolerable.

FIGURE 3: TLF values calculated minus measuredvalues, mean value and standard deviation

CONCLUSIONThe 140-3 standard needs to be revised as regards themeasurement of TLF and the conversion of the SRIresults. This conversion using a reference TLF reducesthe differences between laboratories significantly. Alarge number of field measurements show that, inpractice, the spread of TLF is not so wide as expected,so that the field TLF can be represented by aregression curve. In order to achieve consistencebetween measurements (ISO 140-3) and calculation(EN 12354-1), it is proposed to use the a mean TLFobtained from the field measurements as the referencevalue when the conversion procedure is applied. Ingeneral this will furnish the best results even if thesimplified model is used for calculation.

REFERENCES1. Schmitz, A, et al.: Inter-laboratory test of sound

insulation measurements on heavy walls, part I and II,Journal of Building Acoustics 1999 (6)

2. Fischer, H.-M. et al.: Einheitliches Konzept zurBerücksichtigung des Verlustfaktors bei Messung undBerechnung der Schalldämmung massiver Wände,DAGA 2001 Hamburg, proceedings

dB

959799

101103105107109111113

Frequency f50 100 200 400 800 1600 3150 Hz

TLF

-5-4-3-2-101234

50 100 200 400 800 1600 3150Frequency f

Hz

dB

(2)

Acoustical quality in rooms for mentally challenged people

L. Nijsa, D. van Berloa, D. de Vriesb

a Delft University of Technology, Faculty of Architecture, Building Physics Sectionb Delft University of Technology, Faculty of Applied Sciences, Laboratory of Acoustical Imaging and Sound Control

A research project is going on to improve the acoustical quality in living rooms and work spaces for mentally challenged people.Measurements have been done in living spaces and improvements will be proposed. The aim of the total research project is toimprove existing situations but also to develop guidelines for architects and acoustical engineers for future plans [1].

INTRODUCTION AND METHOD

Long reverberation times and a low speech intelli-gibility occur in living rooms for mentally challengedpeople. The amount of absorbing surface is insufficientand the ratio between challenged people and nursingstaff is kept low, so people live in a “reverberant multi-source environment”. To improve existing situationsand provide for quality in future plans, architecturalguidelines are needed.

The first step in measurements or ray-tracing calcu-lations is to determine the Schroeder integral. Fromthis integral values like the sound pressure level andreverberation time can be calculated. To express theacoustical quality in rooms for mentally challengedpeople we follow Bradley [2], distinguishing “useful”and “detrimental” sound energy, arriving before andbeyond 50 ms. Adding noise energy (NE) from othersound sources as well leads to a variable, commonlydenoted as U50:

+= ∫∫

∞

50

250

0

250 )()(log10 NEdttpdttpU (1)

When NE = 0, the variable U50 is denoted as C50.For one sound source in a diffuse room, the prob-

lem can be encountered from the other side. We departfrom the energy in the direct and reverberant soundfield, writing them respectively as:

20 4/ rWEdir π= and AWErev /)1(4 0 α−= .

W0 represents the source power, α stands for themean absorption coefficient, A represents the total ab-sorbing surface within a (diffuse) room and r stands forthe source receiver distance. Now we find for the use-ful and detrimental energy and for U50:

( )[ ]RTEEE revdiruse /69.0exp1 −−+=

( ) NERTEE revdetr +−= /69.0exp (2a, b, c)

( )detruse EErU /log10)(50 =

RT represents the reverberation time of the room;the noise source is in the diffuse field of the room.

The main advantage of our incorporation of the di-rect sound (left out in most other methods) is that U50

now depends on the source-receiver-distance. This isalso the case in real situations: approaching a speakingperson in a noisy or reverberant room increases thespeech intelligibility.

COMPARING MEASUREMENTS, RAY-TRACING AND THE SIMPLE MODEL

Measurements have been done in four living roomsand two workspaces. Per room two or three sourcepositions were used. The microphone was on a rail of1.50 m long; measurements were taken with 5 cmintervals. Two or three rail positions were used.

The same rooms have been simulated in the ray-tracing model. The main problem is that absorptioncoefficients are required for all surfaces, including fur-niture. From literature absorption coefficients wereestimated for 6 octave bands from 125 to 4000 Hz.RT’s were compared for both methods and α’s (for onesituation) were readjusted with the same factor for allmaterials and all six octave bands. Multiplication fac-tors appear to be surprisingly small as they vary from0.9 to 1.1, depending on the situation. The third stepwas to calculate equations (2) for the same situations.

Figure 1 shows two examples where C50-values arecompared. Some typical results are:• The agreement in mean values between measure-

ments and ray-tracing results is good, but no corre-lation is found for a point to point comparison.

• Variations in measuring results are always biggerthen in ray-tracing results. This has been reportedbefore in literature [3], and is mainly caused bystanding waves in the room. As expected the lowerfrequencies show the highest variations.

• C50-values show the highest variations. The varia-tions in Lp-values and reverberation times (notshown here) are substantially smaller and the agree-ment of mean values is even better.

• C50-values from equations (2) show a fair agree-ment with measurements and ray-tracing in about80% of the cases (represented by figure 1a). In

other cases differences of 3 to 4 dB are found. Thedifferences in the special case of figure 1b can bedecreased from 3 to 1 dB by removing (both inmeasurements and ray-tracing) one table from theroom that acts as a strong reflecting element be-tween source and receiver. Of course, these adjust-ments are impossible in equations (2).

0

2

4

6

8

0 2 4 6 8

measu

rem

ents

[dB

]

125 Hz

250

500

1000

2000

4000

2

4

6

8

10

2 4 6 8 10calculations [dB]

measu

rem

ents

[dB

]

FIGURE 1. C50-comparisons between measurements andcalculations from ray-tracing for two situations. Rectanglesshow the ranges predicted by equations (2).

ACOUSTICAL STANDARDS

It is not easy to find acoustical standards from lit-erature for mentally challenged people. For “normal”ears a value of C50 = +3 to +5 dB is considered as “ex-cellent” for speech intelligibility. The signal-to-noiseratio is often taken as 15 dB. For mentally challengedpeople some literature can be found where a reverbera-tion time in the order of 0.3 to 0.5 is advised, while theS/N-ratio should be 20 dB. However, this demandmeans that U50 should be in the order of 20 dB too, sothe values contradict in a multi-source environment.Other considerations on a basis of the speech transferindex (not given here) lead to our preliminary value of

C50 = U50 = +6 dB for use in the architectural guide-lines we are developing.

ARCHITECTURAL GUIDELINES

The C50-values in figure 1a can be considered as“good” for a living room; the values in figure 1b areeven “excellent”. RT-values are in the order of 0.6 s.However, some of the other living rooms showed re-verberation times as high as 1.1 s and negative C50-values were found. In those cases improvements arestrongly needed.

In single-source situations (where C50-values areused) improvements can be rather simple: adding ab-sorptive materials on the ceiling will be sufficient.However, C50-values turn into U50-values in a multi-source environment. These values can be measured orcalculated in the ray-tracing model by combining twosource positions, but equations (2) also give a goodapproximation assuming that the microphone is in thediffuse field of the “noise source”.

This has been done in the simple method which isthe subject of the Internoise-paper [1] describing thearchitectural guidelines for a first estimation of thetotal amount of absorbing surface. Differences betweenray-tracing and the simple method appear to be small ifabsorbing materials are homogeneously placed alongthe room. If this is not the case, differences may be inthe order of 5 dB.

In a multi-source environment it appears very diffi-cult to reach U50 = +6 dB. Absorbing material on theceiling is not sufficient to provide for a desired valueof A. Absorption on the walls should be used as well.

CONCLUSIONS

The given research shows that a ray-tracing modelcan give a very good approximation of the mean valuesof the acoustical quantities, but deviations in the orderof 4 to 5 dB may occur for specific microphone posi-tions, especially for C50-values.

Simple equations from diffuse field theory are wellsuited for use by architects in the first steps of theirplans to calculate the total amount of absorbing sur-face.

REFERENCES

1. L. Nijs et al., Architectural guidelines to improve theacoustical quality in rooms for mentally challenged peo-ple, Internoise 2001, The Haque, 2001.

2. J.S. Bradley, Speech intelligibility studies in classrooms,J. Acoust. Soc. Am, 1986, 80, 846-854.

3. D. de Vries, J. Baan, Fluctuation of room acoustical pa-rameters at small spatial intervals, Forum Acusticum,Berlin, 1999

The effects of acoustical damping to reduce the structure -borne sound radiation from a double-leaf wall

M. Yairia,b, K. Sakagamia, E. Sakagamia, M. Morimotoa, A. Minemurab and K. Andowb

aEnvironmental Acoustics Laboratory, Faculty of Engineering, Kobe University, 657-8501 Kobe, Japan bKajima Technical Research Institute, 2-19-1 Tobitakyu, Chofu, 182-0036 Tokyo, Japan

Some acoustical damping effects are theoretically investigated to reduce the amplification caused by the mass-air-mass resonance system in the sound radiation of double-leaf walls: an absorbent layer in the cavity, wall surface absorption and a perforated interior wall. As the results, the absorbent layer removes the resonance peak, and efficiently reduces radiated sound power at almost all frequencies. Wall surface absorption is effective around the resonance frequency only, and a perforated interior wall is effective at the resonance, however it can deteriorate the performance at high frequency, depending on its perforation ratio.

INTRODUCTION

Double-leaf walls, composed of a structural wall with an interior wall, cause a large peak in its sound radiation characteristics at low frequencies where the interior wall shows an amplification effect (rather than reduction effect) due to the mass-air-mass resonance [1]. This negative effect must be removed. One of the strategies to reduce the mass-air-mass resonance is introducing an acoustical damping into the system.

In this paper, the effects of the following three types of acoustical damping are studied theoretically: (a) an absorbent layer in the cavity, (b) wall surface absorption, and (c) a perforated interior wall.

THEORY

To consider the effect of an absorbent layer between the two leaves in Fig. 1 (a), it is modeled by a medium of arbitrary propagation constant γ and effective density ρb, written as follows [2]:

( )ωγρ iZb −= , (1)

where Z is the characteristic impedance of the medium and ω is the angular frequency. Both γ and Z can be obtained from an appropriate model, e.g. [3].

The surface absorption of the walls can be included by considering the equivalent particle velocities v’1…v’4 of the medium on the boundaries, which are caused by the acoustic admittances of the surfaces, A1…A4 in Fig. 1 (b), and written as follows:

4,3,2,1,00 ==′ jcApv jjj ρ , (2)

where ρ0 is the air density, c0 is the sound speed in air and p1…p4 are the sound pressures on the surfaces.

In case of a perforated interior panel in Fig. 1 (c), a flow of velocity vf in the hole occurs under the pressure difference ∆p and the velocity vp of the plate itself, and the flow resistance R of the hole is defined as [4]:

( )pf vvpR −∆= . (3) Considering the phase difference in and around the hole, and the viscous damping caused by the friction in the hole, R can be expressed as [5]: FIGURE 1. Geometry of a double-leaf elastic plate of infinite extent, (a) with multi-layered cavity, (b) with

absorptive surfaces characterized by their acoustic admittances and (c) with a perforated interior wall.

A1 A2 A3 A4

p1 p2 p3 p4 vp ∆p

a1

a2

21

1

aa

a

+=σ

z

Structural Wall

Point Force

0

r θ

Interior Wall

γγγγ ρρρρb

(a) (b) (c)

vf

StructuralWall

Interior Wall

z 0

r θ

z 0

r θ

Structural Wall

Interior Wall

Receiving Point

Receiving Point

Receiving Point

Point Force

Point Force

( )2000 28tan dhhkciR σζρ +′−= , (4)

where k0=ω/c0 is the acoustic wave number, h’ is the hole length with open-end correction, d is hole diameter, ζ is the air viscous damping coefficient and σ is the perforation ratio. Since the surface impedance is not uniform due to the holes, an averaged impedance of the panel surface should be taken by following averaged particle velocity v expressed with σ:

( ) pf vvv σσ −+= 1 . (5)

For the three cases above, the radiated sound power

is analyzed by the same procedure as in the previous study [1].

DISCUSSIONS

Figure 2 (a) shows a calculated example of the sound power level with a cavity filled with an absorbent layer. Experimental result, averaged within 1/3-oct band, is also shown for comparison. The theoretical result is in fairly good agreement with the experimental result. Comparing the theoretical results with and without absorbent layer, the peak around 200Hz shift to lower frequency around 80Hz and its value becomes smaller. This is because the mass-air-mass resonance frequency shifts to lower frequencies with decrement of sound speed in the absorbent layer. Moreover, the amplification by the resonance is also damped by the absorption in the cavity.

Figure 2 (b) shows a calculated example of the sound power level with the acoustic admittances of the interior wall’s surfaces. Comparing the results for A1=0.013 and A1…A4=0, no difference is observed at all frequencies. In the case of A2=0.013, the mass-air-mass resonance peak is somewhat damped by absorption in the cavity. The same results are obtained for the structural wall’s surfaces, A3 and A4.

Figure 2 (c) shows a calculated example of the sound power level with the perforated interior wall. The mass-air-mass resonance peak becomes less significant with increasing σ that leads to the reduced value of ∆p. On the other hand, the sound power transmitted through the holes increases with increasing σ at high frequencies. Therefore, its radiation characteristics become close to those of the structural wall alone. When σ exceeds 0.8% in this example, the radiated power becomes larger at high frequencies, whereas the attenuation at the resonance remains almost the same. This fact suggests that there exists an optimal value of σ that maximizes the radiation reduction at the resonance with keeping the negative effect at high frequencies minimum.

ACKNOWLEDGMENTS

The authors thank Ms. Hiroko Fujiwara, Ms. Mina Iizuka and Mr. Yasushi Chiba for their cooperation in this work.

REFERENCES

1. M. Yairi, K. Sakagami, M. Morimoto, A. Minemura and K. Andow, Proc. WESTPRACVII, 1065-1068 (2000).

2. Z. Maekawa and P. Lord, Environmental and architectural acoustics, E&FN Spon, London, 1994, pp328-329.

3. Y. Miki, J. Acoust. Soc. Jpn. (E) 11, 1, 19-24 (1990). 4. A. D. Pierce, Acoustics, McGraw-Hill, NY, 1981, pp. 146-148. 5. D. Takahashi, Appl. Acoust. 51, 71-84 (1997).

FIGURE 2. Effects of thee types of acoustical damping in a double-leaf wall, with an interior gypsum board 4.5kg/m2 and a concrete wall 460 kg/m2: (a) The cavity 0.06m deep filled with absorptive layer. (b) Acoustic admittances of the surfaces with an air cavity 0.03m deep. (c) Perforated interior wall with the holes of 4mm diameter. An air cavity is 0.03m.

σ=1.5%

σ=0.8%

σ=0%(No perforation)

σ=0.4%

A1=0.013 and A1…A4=0

A2=0.013

R=10kPa s/m2

Air only

Experimental data

(b)

(c)

(a)

Subjective Evaluation of Impact Sound Transmissionthrough Floor Structures

E. Nilsson, P. Hammer

Department of Engineering Acoustics, LTH, Lund University, P. O. Box 118, SE-221 00 Lund, Sweden

The impact sound from eight different floor structures has been judged in paired comparison tests. The correlationbetween preference values given by the listening test and different measures of impact sound has been calculated.As a result of the investigation it follows that the loudness measure according to ISO 532B gives the highestcorrelation to the preference values given by the listening test. Thereby it is concluded that loudness is a suitablemeasure for evaluating the subjective impression of impact sound. Further it works both for lightweight and heavyfloor constructions. Comparable with loudness the commonly used weighted sound pressure level Ln,w (EN ISO717/2) yields considerably lower correlation to the preference values. This is especially emphasised for lightweightconstructions.

INTRODUCTION Building multi-storey wooden house using light-weight wall and floor structures has increased the needof relevant methods and measures for evaluation of theacoustical conditions. Especially concerning impactsound insulation the lack of correspondence betweenstandardised evaluation methods and subjectiveimpression has been observed [1, 2]. The purpose of this project is to investigate thecorrelation between commonly used measures ofimpact sound and subjective judgements from listeningtest. Further to investigate if psychoacoustics measuresor combination of these can be used to improve thiscorrelation. Both impact sounds generated by normalwalking and by the standardised tapping machine isincluded in the analysis.

METHOD The recordings of the impact sound were performedas a two-channel measurement using a dummy head.Two sources of impact sound generators were used, amale walker with a weight of 95 kg and a femalewalker with a weight of 65 kg. In the listening test thesubjects were listening to the recordings usingheadphones. The dummy head recordings wereperformed in a laboratory for impact soundmeasurements on floor constructions. However, theabsorption in the receiving room was somewhatincreased to correspond to a normal living room. Themeanvalue of the reverberation time in the frequencyregion 250 to 2000 Hz was about 0.5 seconds. A total of eight floor constructions were included inthe study. Five of these were wooden structures and

three were based on a concrete slab. The floors aredescribed in Table 1.

Table 1. Floors included in the test

1. 22 mm chipboard, 250x95 wooden joist c600, 2x95mineral wool 20 kg/m3, 2x13 plasterboard

2. 22 mm chipboard, 25 mm mineral wool, 22 mmchipboard, 250x95 wooden joist c600, 2x95 mineralwool 20 kg/m3, 2x13 plasterboard

3. 22 mm chipboard, 25 mm mineral wool, 22 mmchipboard, 250x95 wooden joist c600, 2x95 mineralwool 20 kg/m3, resilient channel, 2x13 plasterboard

4. 350x350x25 concrete plates, 22 mm chipboard, 25 mmmineral wool, 22 mm chipboard, 250x95 wooden joistc600, 2x95 mineral wool 20 kg/m3, resilient channel,2x13 plasterboard

5. 22 mm chipboard, 250x95 wooden joist c600, 2x95mineral wool 20 kg/m3, resilient channel, 2x13plasterboard

6. 160 mm concrete7. 22 mm chipboard, 25 mm mineral wool, 160 mm

concrete8. 22 mm chipboard, 3 mm foam, 160 mm concrete

The listening test comprises a paired-comparison testof eight floor constructions. Two floor constructions ata time were compared and the subjects were asked tojudge, which of them that sounds most annoying (mostunpleasant and irritating). Ties were allowed in thetest. Thirteen subjects listening to male walk (95 kg)and ten to female walk (65 kg). All the subjects werenormal hearing. The correlation between subjective judgements andobjective measurements was investigated. By means ofthe correlation factor the suitability of the parameters

in Table 2 as tools to render the subjective impressionof the impact sound is tested.

Table 2. Objective parameters

Parameter DescriptionLn,w Evaluation according to EN20 ISO 717/2.

Ln,w,8 As above but with limited maximum deviation8 dB between reference curve and measuredcurve.

LB As EN 20 ISO 717/2 but with a reference curvewith a slope of 1 dB octave band starting at 50Hz and ending at 1000 Hz [1].

Leq,A,F A-weighted equivalent sound pressure with livefootstep as exciter.

Leq,A,H A-weighted equivalent sound pressure with theISO tapping machine as exciter.

Leq,C,H C-weighted equivalent sound pressure with theISO tapping machine as exciter.

NF Loudness* according to ISO 532 B with livefootstep as exciter.

NH Loudness according to ISO 532 B with tappingmachine as exciter.

NF, SF A combination parameter of loudness andsharpness* with weighting factors given by alinear regression to the ranking values. Livefootstep sound as exciter.

NH, SH As above but with tapping machine as exciter

* These measures are discussed in [4]

RESULTS

The results from the listening test have been analysedusing a model by Rao and Kupper [3]. As a result fromthe model rating values together with a 95 %confidence intervals are calculated. The rating valuesand confidence intervals for the eight floorconstructions are shown in Figure 1.

Figure 1. Rating values with 95% confidence interval. Bothmale and female walk is included in the analysis.

Floor number 4 and 7 are judged as less annoying incomparison to the other floor constructions. Floornumber 2, 3, 5, 6 and 8 are judged as almost equallyannoying and floor number 1 is judged as the mostannoying floor construction. The correlation coefficients between the objectiveparameters according to Table 1 and the rating valuesare given in Table 3.

Table 3. Correlation between rating values andobjective parameters.

Correlation coefficientPara-meter Female walk Male walk Both female an

male walkLn,w -0.90 -0.77 -0.84

Ln,w,8 -0.87 -0.74 -0.81LB -0.90 -0.90 -0.92

Leq,A,F -0.93 -0.90 -Leq,A,H -0.91 -0.76 -0.84Leq,C,H -0.86 -0.89 -0.89

NF -0.92 -0.91 -NH -0.92 -0.93 -0.94

f(NF,SF) 0.95 0.91 -f(NH,SH) - - 0.94

From Table 3 it appears that loudness according to ISO532 B give the highest correlation to the preferencevalues. It is even so that the loudness values given bythe tapping machine yields higher correlation thanloudness values given by real footstep. The reason forthis is not investigated but it seems possible thatbackground noise will influence the results for the casewith real footstep. The lowest correlation is obtainedfor the parameters Ln,w,8, Ln,w and Leq,A,H. It is noticeable that the combination parameterincluding both loudness and sharpness increase thecorrelation coefficient for the female walk but not forthe male walk. Thus, including the sharpness parameterthe relatively larger amount of high frequency contentin the female walk is rendered in the combinationparameter.

REFERENCES

1. K. Bodlund, Alternative reference curves for evaluation of theimpact sound insulation between buildings. JSV, 1985, 102(3),pp 381-402

2. P. Hammer and E. Nilsson, On subjective grading of impactsound transmission through lightweight floor structures. Inter-Noise 97, Volume II, Budapest, Hungary

3. P.V. Rao and L.L. Kupper, Ties in paired comparisonexperiments: A generalisation of the Bradley-Terry model.

4. E. Nilsson and P. Hammer, Subjective evaluation of impactsound transmission through floor structures. Report TVBA-3103,Engineering Acoustics,LTH, Lund University, Sweden.

0,001

0,01

0,1

1

0 1 2 3 4 5 6 7 8 9

Floors

Rat

ing

valu

es

On the mechanics of gypsum plate systems

Jonas Brunskog, Per Hammer

Division of Engineering Acoustics, Lund University, P.O. Box 118, SE-223 63 Lund, Sweden

Lightweight walls and ceilings are commonly used in buildings. It is often praxis to use two layers of gypsum plates. A gypsumplate consists of a core of gypsum surrounded by thin sheets of paper on each side. The paper sheets are working asreinforcement. In order to describe this system properly, the plate has to be seen as a sandwich element. When two plates are puttogether, air will be trapped in a narrow layer between the plates. This air layer will influence the coupling and dissipation of thesystem. An elaborate theory of this coupling is indicated in the paper, whereas a simplified theory is given in a separate paper.The influence of screws is then discussed.

INTRODUCTION

In the building industry, as well as in other branches,lightweight techniques become increasingly important.For lightweight walls and ceilings in buildings, it isoften praxis to use two layers of gypsum plates.Gypsum plates consist of a core of gypsum surroundedby thin sheets of paper on each side. Due to the papersheets, the plate has to be seen as a sandwich element.When two plates are put together, air will be trapped ina narrow layer between the plates. Depending on thewidth of the air layer, it will influence the couplingand dissipation of the system. An elaborate theory ofthis coupling is indicated in the present paper, whereasa simplified theory is given in a separate paper [1].The influence of screws is also discussed.

THE GYPSUM PLATE

The paper sheet will make the influence of the sheardeformation in the gypsum plate more important thanin ordinary Kirchof plate theory, so that sandwichtheory has to be used. Due to the predominant sheerdeformation the results is likely to be similar to thetheories presented by Mindlin [2]. A variationalapproach is used, and some of the strain energies aretaken from Mindlin. The geometry and displacementsare shown in figure 1. The displacement w and therotations θ1 and θ2 are also indicated.

d

h

θ

x1

x3

x2

w

θd/2 FIGURE 1. Geometry of the sandwich plate.

ξR p- τ21-

p+ τ21+

τ23-

τ23+

w x1

x3

x2

FIGURE 2. Pressure and shear acting on the sandwichplate.

With the aid of Hamilton’s principle, and after somemanipulations, the governing equations for thedisplacement w and rotations θ1 and θ2 is found to be

( ) ( )( )( )

( ) ( )( )( )

( ) pwdiwdG

dJixwdG

xDxhdE

dJixwdG

xDxhdE

R

x

R

−=++∇′−=++∂∂′−

−∇−+∂∂++∂∂

−=++∂∂′−−∇−+∂∂++∂∂

ρωφτβωθ

θνφνθτβωθ

θνφνθ

2

23222

22

2222

22

2111

12

1211

22

2

2112

2

2112

where ER is the Young’s modulus in the paper sheet,G’=κ2G and

( ) yxdJEID yx ∂∂+∂∂==−= θθφρν ,12,1 32

If ER=0 and the external shear forces are zero, thisexpression equals Mindlin’s results. Thus, the flexuralsandwich equations for a gypsum plate can be seen asa modified Mindlin plate.

THE NARROW AIR LAYER

A thin layer of air will be trapped between the plates.

x1

x2 x3

FIGURE 3. Geometry of trapped air.

The field variables are; pressure p, mass density δ,temperature τ, entropy per unit mass σ and the velocityu=(u1, u2, u3) without any constant part. The Stokes-Navier equation, the continuity of mass flow, thecontinuity of heat flow, and the two equations of statefor a perfect gas is used to couple the field variables.However, most of the field variables can be excluded,so that the solution will be given with u and τ asprimary field variables. Achenbach [3] has given a general approach fortreating Lamb waves in homogenous plates asthickness vibrations superimposed on a membranecarrier wave. This approach is here used for the termo-viscous layer of trapped air. Following Achenbach, we chose the following formfor the velocity

( ) ( )

( ) ( ) .,

2,1,,1

2133

213

ti

ti

ii

exxxzu

iex

xxxv

ku

ω

ω

ϕ∂

ϕ∂

=

==

Here k is a wavenumber-like quantity, i.e. itsdimension is 1/length. The function ϕ isdimensionless. In the same way, the temperature isassumed to be of the form

( ) ( ) tiexxxyt ωϕτ 213 ,),( =x

The assumption that the same x1-x2-dependence is usedfor all field variables can be made probable as it isequivalent to assuming a e-ikx dependence. We willseek steady-state solutions, and we will omit exp(iωt).After some manipulations, the solution of the field inthe x3 direction may be written as a combination ofthree modes a, b and c

( )( )( )

333231

3

3

3xxx eCeBeA

xy

xz

xvκκκ ±

±±

±±

± ++=

cba

where the constants A±, B± and C± are to bedetermined by the boundary conditions. The x1-x2-dependence is described by the function ϕ( x1,x2),generated from the membrane equation

022 =+∇ ϕϕ k ,

and where k is to be a solution of a dispersion relationsyielding from the boundary conditions.

THE COUPLED SYSTEM

Let the boundary conditions for the termo-viscousfluid be

( ) ( )

( ) ( )

( ) ( )

( ) ( )

( ) ( )( ) ( )

( ) .0,,

,,,,

,,,,

,,2

,,

,,2

,,

,,2

,,

,,2

,,

21

21)2(

211

21)1(

213

21)2(

2

)1(

212

21)1(

2

)1(

212

21)2(

1

)1(

212

21)1(

1

)1(

211

=±=−

=±

−=−

=

−=−

=

hxx

xxwihxxu

xxwihxxu

xxd

ihxxu

xxd

ihxxu

xxd

ihxxu

xxd

ihxxu

τωω

θω

θω

θω

θω

so that two plates (superscript 1 and 2) is coupled viathe termo-viscous fluid. The plates are assumed tohave zero temperature fluctuation. Moreover, the x1-x2-dependence will be the same for all layers in thesystem, which reduces the size of the problem andguarantees that it is solvable.

SCREWS AND NAILS

The screws and nails are included by means ofintroducing a periodic array of boundary conditionsthat couple the different fields. This boundarycondition is first seen as force

( ) ( )2211, nlxnlxFn m

mn −−∑ ∑∞

−∞=

∞

−∞=δδ

applied to the equations. In the next step these forcesare solved so that the boundary conditions is fulfilled.

CONCLUSIONS

It is possible to find a elaborate theory for themechanics of gypsum plates that includes sandwichtheory for the plates with termo-viscous theory for thetrapped air.

REFERENCES 1 Lindblad, S. G., and Brunskog, J., ‘On the air-gap betweennarrow plates’, in Proceedings of 17th InternationalCongress on Acoustics, 20012 Mindlin, R D., J. Applied Mechanics, 18, 31-38 (1951).3 Achenbach, J. D., J. Acoust. Soc. Am. 103 (5), 2283-2286(1998)

On the air-gap between narrow plates

Sven G. Lindblad, Jonas Brunskog

Division of Engineering Acoustics, Lund University, P.O. Box 118, SE-221 00 Lund, Sweden

It has been reported in experimental investigations that the air-gap between gypsum plates can have an influence on the soundtransmission of lightweight walls. This small influence can be explained as a resonance. In theoretical literature one can find theinfluence on the energy dissipation of a narrow air layer behind a plate, but not on the transmission. The aim of this paper is todescribe the influence of a thin air layer on the sound transmission. The paper investigates this phenomenon by means of asimple theory and gives a historical background to the theory. The narrow air layer is described by means of wave motion in theplane of the plates, i.e. not as a locally reacting spring. Viscous effects are included by means of a parabolic displacementassumption.

INTRODUCTION

It is common practice to use plates close togetherwith a layer of air trapped in-between in buildingsystems. Examples of this is windows, gypsum platesin lightweight walls and floors and in parquet types offloor covering. It has been reported [1] that the airtrapped in the gap between plates can have aninfluence on the sound transmission. This smallinfluence can be explained as a resonance. Intheoretical literature one can find the influence on theenergy dissipation of a narrow air layer behind a plate,but not on the transmission. The aim of this paper is todescribe the influence of a thin air layer on the soundtransmission.

AIR-GAP RESONANCE

A narrow air gap between thin plates in soundinsulating structures gives rise to a mass-compliance-mass (m’’1-C-m’’2) resonance at a fairly highfrequency

( ) ( ) Cmmf 12

11

10 2 −−− ′′+′′= π (1)

Normally this is avoided by various reasons, but insome cases the design gives say 0.5 mm airspace. Atraffic noise sealed window with three panes is anexample. In this case the resonance for the 0.5 mm air-gap would be about 1 kHz. Measurements showedalmost no effect in the TL-curve (Transmission Loss)of this high resonance caused by the narrow airspace.One reason for this could be losses in the complianceleading to a complex compliance Cc. This type oflosses has been considered many times in the past, e.g.[2, 3], in order to introduce extra damping. The resistance has great effect at grazing angels ofincidence, where Cc would tend to zero without it and

when resonance occurs, especially when the air spaceis less than 0.5 mm. Results according to the abovereasoning were given by the senior author at ICA1974 in London [4]. Recently Warnock [1] has pointed out that in othersituations the narrow space resonance can be seen inthe TL. In a large measurement series of lightweightjoist floors the TL for floors with a single layer ofgypsum board were compared with TL for two sheetsof the material. When two plates are used the twosheets of material is not perfectly in contact and an air-layer is trapped. The difference in TL for the two casesshowed a clear dip at 1 kHz, and Warnock estimatedthe trapped air to be 1 mm thick.

VISCOUS EFFECTS

The losses in the compliance are due to viscosity. Inthe thickness direction wave motion is neglected asthe space is narrow. On the other hand there is wavemotion along the space leading to alternatingstreaming guided by a flow resistance σ. This leads toa modified, complex compliance Cc.

ρωσβ

ρ ic

dCc +

−=1

sin1

2

2 (2)

where ρ is density, c sound velocity, β is the angle ofincidence, ω is the angular frequency, and d is thewidth of the air layer. The resistance σ follows fromassuming a parabolic flow profile in the narrow space,σ=12ν/d2, with the dynamic viscosity normally takenas ν = 18⋅10-6 kg/m3s. The laterally produced losses are of relaxation type,yielding approximately σ/ωρ=1 as optimal, see figure1. If on the other hand σ increases to infinity C isagain purely real, and the same occurs for σ=0. In aspecific design situation it is important that the region

of optimal imaginary part coincides with theresonance.

ωρ/σ -3 -1 1 10 10 10 10 3

-0.4

-0.3

-0.2

Im{Cc}ρc2/d -0.1

0

FIGURE 1. Imaginary part of the complex compliance Cc

as a function of a frequency parameter, when sinβ=1.

TRANSMISSION LOSS

It is now the aim to find the transmission loss for thedouble plate system. London [5] was the first todescribe the transmission loss for the double platesystem. A system consisting of two plates coupled viacompliance is e.g. given in Cremer and Heckl [6].Later Kropp and Rebillard [7] studied some aspects ofairborne sound insulation of double walls, and theircalculations are in much followed here. If we includethe viscous air-layer the displacements of the plates aregiven as

( )( ) ,

,

222

1222

112

2111

tc

ric

pwmKwwwB

ppwmKwwwB

−=′′−−+∆∆

+=′′−−+∆∆

ωω

(3)

where pi, pr and pt are the incident-, the reflected- andthe transmitted pressures respectively, and Kc=1/Cc.For z=0, we assume an incoming wave in the x-direction with the wave number kx=ksinβ. On bothsides of the construction the displacement in air andthe displacement of the plate have to be identical.Additionally, the impedance of the wave field of bothsides gives the ratio between pressure and velocity as¨

10 2 10 3

5

15

25

35 TL(β) [dB]

d=1e-6 d=1e-4 d=5e-4 d=1e-3 d=2e-3

FIGURE 2. Transmission loss for two gypsum plates withvarious width of the air-gap. Theoretical values. Angel ofincidence 45°.

( ) βρωω cos,ˆ,ˆ 02010 cZwiZpwiZpp tri =−==− (4)

Using a spatial stiffness Si,=Bikx4-ω2mi’’, and after

some algebra, we have an expression for thetransmission, used to calculate figure 2.

( )( )( ) 2

0201

02

ccc

c

KiZKSiZKS

ZK

−++++=∆

∆=

ωω

ωβτ (5)

ANALYSIS OF MEASUREMENTS

In order to investigate if the air-gap resonance couldbe seen also in the measurements preformed at thedivision of Engineering Acoustics in Lund. A numberof lightweight walls with both one and two layers ofgypsum board were use [8]. In figure 3 the differencein TL is shown. A minimum is seen around 1.5 kHz.

frequency [Hz] 63 125 250 500 1k 2k

-1 0 1 2 3 4 5 6 7 8 9

∆ TL [dB]

FIGURE 3. Difference between single sheet and doublesheets of gypsum board, the solid line is the mean.

CONCLUSIONS

The air-gap resonance can have influence of thetransmission loss, and is therefore an important designquantity.

REFERENCES 1 Warnock, A.C.C. Airborne and Impact Sound Insulation ofJoist Floor Systems: A Collection of Data, NRCC-44210,National Research Council Canada, Monteral, 20002 Cremer, L. and Müller H. A., Principles and Applicationsof Room Acoustics, Applied Science, London, 19823 Trochidis, A., Acoustica, 51 (4), 201-211 (1982)4 Lindblad, S. “The influence of radiation efficiency andinterspace wave on the transmission through multiple leafpartions”, in Proceedings of 8th International Congress onAcoustics, 1974, p. 673 .5 London, A. J., Acoust. Soc. Am., 22 (2), 270-9 (1950)6 Cremer, L. and Heckl, M. Structure-borne sound,Springer-Verlag, Berlin, 1973, second edition 1988.7 KROPP, W. and REBILLARD, E., Acustica, 85, 707-720(1999)8 Hammer, P. and Nilsson, E. Isoleringens inverkan påljudisoleringen för lätta väggar och bjälklag, Report TVBA-3092, Lund University, Lund, 1996 (in Swedish)

Flanking Transmission in Lightweight BuildingsL. G. Sjökvist and P. Hammer

Department of Engineering Acoustics, Lund University, John Ericssons väg 1, 223 63 Lund, Sweden

An experimental study was recently performed for different kinds of flanking transmission paths. In the experiments, the lightweightfloor beams orientation were tested. A prediction was made according to the EN 123454-1. The study shows that EN 12354-1 predictsthe transmission loss lower than all tested settings. Result shows that floor beams orientation influence the sound transmission.

EXPERIMENTAL SET-UP

Experiments were made, similar to the ISO 140 stan-dard, to measure the sound transmission loss between tworooms including flanking transmission. A double wall di-vided the two rooms, and eight different conditions weremeasured.

In the first four settings the floor was made with beamsperpendicular to the wall. In the cavity at floor level threeconnections of different materials were attached and onemeasurement was made with no connection between thefloors. The different materials tested were 22 mm Fibre-board, 1 mm Steel nail joints and 0.5 mm Steel plates.The board was placed continuously in the cavity whilethe steel was placed centred at 600 mm distance. Thesteel plates dimension were 140 mm× 160 mm. Thetests were repeated for floors with beams parallel to thewall.

The floors were 4.00 m× 4.70 m in the receiver roomand 4.00 m× 3.50 in the sender room. The wall was 4.00m× 2.22 m. The intersection was constructed as shownin figure 1 at Lund University in February 2001.

160

22 FIBERBOARD 215 x 45 JOIST C.600/

MINERAL WOOL 22 x 95 JOIST C.600

10 FIBERBOARD

2x13 GYPSUM 95x45 JOIST c.600 200 Air Gap 95x45 JOIST c.600/

MINERAL WOOL 2X13 GYPSUM

4000

2215

3500

CONNECTOR

FIGURE 1. Test construction

TEST RESULTS

Result from measures is shown in figure 2.

63 125 250 500 1000 2000 400020

30

40

50

60

70

80

Frequency

Tra

nsm

issi

on lo

ss [d

B]

FIGURE 2. Test results for parallel floor beams (circled points)and orthogonal floor beams (boxed points) with no connection(homogeneous lines), fibreboard (dashed lines), 1 mm steel nailjoints (dotted lines) and 0.5 mm steel plates (dash-dotted lines)

COMPARISON WITH EN 12354-1

EN 12354-1[1] standard was followed as far as pos-sible in predicting the transmission loss. Notions herefollow the Annex A in EN 12354-1.

The radiation factor (σ) was calculated from

σ =lλ

Z π

0

(

sin(√

fc/ f −cosαkl/2)

(√

fc/ f −cosα)kl/2

)2

dα (1)

where l is the shortest length of the element,λ is thewavelength,fc is critical frequency,f is frequency,α isthe radiation angle andk is the wave number.

For calculation of flanking transmission formulationsfor Ki j andDv,i j , in the EN 12354-1 pp. 49 and 15, wereused. The result in figure 2 for no flanking connectionwas used asRwall and asRf loor a floor construction mea-surement from Homb, Wheem and Strom [2] was used.

Simplification was made so the floor areas on both sideswere 14 m2.

For flanking transmission, the standard recommendscalculation method for some typical cases. This makesinventions for flanking part invisible, as can be seen infigure 3 where some measured situations are shown to-gether with the EN 12354-1 calculation. As one can seethe calculation method is here on the high radiation sidefrom the spread of measured results and that is to be ex-pected while other solutions are made for better acousticperformance.

125 250 500 1000 200030

35

40

45

50

55

60

65

70

Frequency

Tra

nsm

issi

on lo

ss [d

B]

FIGURE 3. Calculation with EN 12354-1 (with boxed points)together with some typical flanking parts (with circled points).

INSERTION LOSS

By insertion loss we mean transmission loss differencebetween the system without flanking connection and thesystem with flanking connection.

In the low frequency range the insertion loss is depen-dent of the floor orientation. In the system with floorbeams parallel to the separating wall, some frequencieshave considerably higher transmission than others. Thesefrequencies also appear too regularly to have happenedby chance. When floor beams are orthogonal to the sep-arating wall we observed lower transmission than for theparallel case within the lower frequency range.

For the negative insertion loss at lower frequencies,one might suspect that the floor act like a Helmholtzresonator. The damping will then be larger in the casewhere more energy is transmitted directly between thetwo floors. This can explain why positive insertion loss isfound in many cases.

In the higher frequency range, one can note that thetransmission loss is dependent on the type of connectorat the junction. This is especially obvious in the case

with fibreboard connector. This indicates that continu-ous plates of the same material do not only have highertransmission, it is also less dependent from the floor beamorientation in this frequency range.

63 125 250 500 1000 2000 4000−25

−20

−15

−10

−5

0

5

FrequencyIn

sert

ion

loss

[dB

]

FIGURE 4. Insertion loss for parallel floors (with circledpoints) and orthogonal floors (with boxed points) with fibre-board (homogeneous lines), 0.5 mm steel-plates (dotted lines)and 1 mm steel nail joints(dashed lines) .

REMARKS

The EN 12354-1 standard overestimates the transmis-sion in lightweight buildings by 1 to 8 dB.

The orientation of the floor beams is important fortransmission in the low frequency range.

Continuous floor plate as connector transmits the sameamount of energy at high frequency independent of floorbeams orientation. This connector also has the best agree-ment with the EN 12354-1 standard.

REFERENCES

1. EN 12354-1:2000, English versionBrussels, CEN, 2000

2. A. Homb, S. Hveem and S.Strøm,Lydisolerende konstruk-sjoner Datasamling og beregningsmetode, Norway, Norgesbyggforskningsinstitutt, 1993 pp. 74.

3. L. Cremer, M. Heckl and E.E. Hungar,Structure-BorneSound, Second edition, Germany ,Springer Verlag , 1988,pp. 101.

Empirical verification of the prediction model designed to estimate the flanking transmission in building

B. Szudrowicz, A. Izewska

Department of Acoustics, Building Research Institute, Ksawerow Str. 21, 02-656 Warsaw, Poland

Measurement verification of the calculation method specified in EN-12354-1:2000 [1] of determining apparent sound reduction index of building elements relies in comparing the results of the measurements obtained according to EN ISO 140-6. This verification can apply both to the detailed method (calculations and measurements conducted in the respective frequency ranges) and the simplified method (calculations are made with respect to the weighted indices Rw; the results are confronted with the indices determined on the basis of the measurements). Partial verification of the calculation method can be conducted through determining, with the use of calculations and measurements, the resultant weighted flanking sound reduction index of flanking paths Ff and Df. Analogically as in case of full verification, the measurements can be based on the detailed or the simplified method.

ASSUMED METHOD OF VERIFICATION THROUGH

MEASUREMENTS We have verified the calculation method for the resultant weighted flanking sound reduction index of flanking paths Ff and Df. The calculations were made with the use of the simplified method. This value was determined in accordance with the following equation (having assumed the radiation factor of 1=σ ).

f

sfvDfFf S

SLLR lg105,271)( ++−=+ (1)

where: L1 - is the average sound pressure level in

the source room, in dB Lfv - is the velocity level (at Vo = 10-9 m/s)

determined on the flanking partition in the receiving room, in dB

Ss - is the area of the separating element area, in m2

Sp - is the area of the flanking partition, in m2 Denoting the sum of the summands in the equation (1), dependable on the construction, as R(Ff+Df)o,m, we obtain the following:

f

smoDfFfmDfFf S

SRR lg10,)()( += ++ (2)

The equation (2) can be applied to the weighted indices R(Ff+Df)o,w. In the simplified method we have the relation:

f

scalwoFDfFfcalwDfFf l

SRR lg10,,)(,)( += ++ (3)

Therefore, the comparison of the measurement and the calculation results can be reduced to verification of the following patterns:

pcalwoDfFfmowDfFf lRR lg10,,)(,,)( += ++ (4)

where lP – is the dimension of the flanking element perpendicular to the edge lf. If the vertically flanking transmission is taken into account then lp = h (where h – is the height of the room). PROGRAM AND CONDITIONS OF THE

EXPERIMENT The tests of the vertically oriented flanking transmission (from bottom to top) were performed in large-panel buildings of varying thickness of walls (14 cm and 15 cm) and floors (14 cm and 16 cm) with floating floors. The height of the tested rooms was h = 2.5 m. Diagrams of the tested junction and the measurement and calculation results have been matched in Table 1. On the basis of the measurements, R(Ff+Df) was determined for the 1/3–octave bands within the 100 – 3150 Hz frequency ranges. For each of the situations presented in Table 1, a number of measurements were conducted in one building (the number of measurements is given in the table), followed by determining the average values and the standard deviation. Standard deviation in the low frequency band is considerable, σ(n-1) ≈ 3 ÷ 5 dB, while in the middle and high frequency bands it decreases to the value of σ(n-1) ≈ 1 ÷ 2 dB. The weighted index given in the table was determined on the basis of average R(Ff+Df)o measurement results obtained within one building.

The KFf and KDf values were determined with the use of the formulas given in the EN standard. The weighted sound reduction indexes of the elements were derived from the law of mass (determined pursuant to the tests performed by the Acoustics Department of the ITB) in the following form:

6,24lg9,30 ' −= mRw (5)

Table 1 implies that the following differences between the calculation results and the test results were obtained:

no of cases ∆Rw construction junction:

24 19

2.0 to 2.5 dB - 1.0 to - 0.5 dB

non-construction junction:

15 - 2.4 to –1.0 dB

FINAL REMARKS

1. The described method of empirical verification of

calculation method can be used only in case of massive constructions, for which σ = 1. However, this method is a greatly advantageous. It allows to eliminate altogether the influence of possible indirect sound transmission (e.g. through leaks) on the measurement results, which can be of great

importance with the above given statements of the measurement and calculation results.

2. The calculations for the construction junction yielded results different from the measurement results by the range – 1 to + 2,5 dB. Taking into account that the calculations were performed with the use of the simplified method, this result can be deemed satisfactory. In case of non-construction junction, the fact that the measurement results are better than the calculation results can be the effect of less rigid connection between the elements in the junction.

3. The differences between the measurement and the calculation res ults are to a large extent dependable on the input results for the calculations. Therefore, the comparison results cannot be treated as direct evaluation of the calculation method. If calculation of the Rw was performed pursuant to other mass law (e.g. according to the relation F given in the figure B.2 in the standard EN-12354 – 1:2000), there would be much greater differences.

REFERENCES

1. EN 12354 – 1:2000, Building Acoustics – Estimation of acoustic

performance of buildings from performance of products. Part 1: Airborne sound insulation between rooms

Table 1. Comparison the measurement and calculation results of flanking transmissiom

Situation within the building Measurement

results

No.

Diagram h1

[cm]

h2

[cm]

n*) R(Ff+Df)o,w,m

[dB]

Calculation results

R(Ff+Df)o,w,cal +10lg(lp)

[dB]

Calculation-measurement

difference

[dB]

1 14 16 11 62 64.0 + 2

2 13 61 63.5 + 2.5

3 15 14

7 64 63.5 - 0.5

4

Construction joint

14 14 12 64 63.1 - 0.9

5 8 (gypsum)

14 7 55 56.6 - 1.6

6

Non-construction joint

5 (concrete)

14 8 60 57.6 - 2.4

*) n – is the number of measurements in the given building

Highly sound-insulating wooden floor systemwith granular filling

M. Walk and B. Keller

Chair of Physics of Buildings, Swiss Federal Institute of Technology (ETH), 8093 Zurich, Switzerland

The goal of this work is the development of a wooden floor system reaching very good sound insulation parameters even withouta large number of different layers. A new approach is proposed with a walking surface floating freely on a layer of granularmaterial, which is known to have a strong internal friction. The floor system presented here is characterised by a high degree ofprefabrication, low production costs and simple installation at the construction site. It attains a standard impact noise level of60 dB, with carpet even clearly below 50 dB. The main characteristics and their impact are presented.

INTRODUCTION

In recent years, also due to the pledge forsustainable development, there has been an increasingdemand for wood as a renewable building material,also in a multi-storey context. One of the majorobstacles is the lack of cheap and reliable solutions forsufficient sound insulation of timber floors. A numberof research projects [1] have tackled the problem byadding several additional layers above and below thecore floor element. Only little effort has been made toimprove the core element itself. This paperdemonstrates such a basic improvement to be possibleand presents a floor system that has been developedusing this approach. The main focus of this researchhas been on impact sound transmission.

THEORETICAL CONSIDERATIONS

In a first step, the most important parameters forimpact sound transmission through wooden floors wereidentified using the well-established theory of vibrationof plates [2] and the given material properties of woodand its composites [3]. Two crucial requirements couldbe deduced:

• the core floor has to consist of at least two layerswith a minimum of resonance frequency overlaps;

• since lightweight materials like wood cannot inhibitsound transmission simply by their mass, adifferent physical basis must be found to achieve agood sound insulation. One possibility is theattenuation of the sound energy by internal frictionand thus its conversion into heat.

IMPLEMENTATION INTOFLOOR DESIGN

Granular materials are well known for their highinternal friction [4], which is due to friction at thecontact surfaces between neighbouring grains. In woodindustry, granular material like sawdust accumulates aswaste product, and is thus cheaply available.

Based on these facts, a concept of a sandwichstructure with granular filling was designed (Fig. 1).To make sure the sound has to pass through thegranular layer on its way through the structure, thewalking surface must be floating on the granular fillingwithout any rigid or elastic connection to the load-bearing structure. The latter, a wood composite platestiffened by joists in longitudinal direction accordingto the static needs, is designed like a trough which actsas a container for the granular material. Also thewalking surface is stiffened by ribs in both longitudinaland transverse direction, in order to minimise itsmotion. By suitably dimensioning the stiffeners ofload-bearing plate and walking surface, commonresonances can be avoided.

FIGURE 1. Sketch of the basic concept of the floor design.(Front termination left out for clarity.)

EXPERIMENTAL INVESTIGATIONS

Different versions of the proposed floor elementwere subject to experimental investigations on a full-size model. The tested variants were characterised bydifferent thickness of the granular layer, differentgranular materials (sawdust, sand) and varyingstiffness of the walking surface. Vibrations wereexcited by standard tapping machine, impulse hammeror electromagnetic shaker. To better understand thevibrational behaviour of the floor, local accelerationswere measured, and evaluated using modal analysistechniques [5].

Based upon the measurement results, an optimisedfloor element was designed which also fulfils the majorrequirements for its usability in real buildings: highdegree of prefabrication (including granular filling),good transportability, easy assembly at theconstruction site and competitive production costs.

A prototype of the optimised element wassubsequently tested under standard conditionsaccording to ISO140.

RESULTS

Fig. 2 summarizes the standardised impact soundlevels L’n,w obtained with some of the variants undertest and relates them to a typical value for a massivefloor. The prefabricated element reaches 60 dB andthus performs 10 dB better than a concrete floor, buthas only one fourth of its mass.

The influence of the distance of stiffening ribsattached to the walking surface has also beeninvestigated by modal analysis. As it is clearly seenfrom Fig. 3, there are strong plate resonances governed

by the size of the fields between the ribs. Theirfrequencies can therefore be tuned by simply adjustingthe rib distance. These resonances were found to beinnoxious if they are tuned to occur above600…800 Hz. This can be reached by rib distances of30…40 cm with the timber plate used here for thewalking surface.

It is also remarkable that the use of a carpet on theproposed floor element reduces the standard impactsound level by 12 dB, a value which is much morecommon for massive than for lightweight floors.

The sound transmission loss of the prefabricatedelement for airborne noise was determined as 52 dB bystandard measurement.

ACKNOWLEDGMENTS

The current research project was funded by theCommission for Technology and Innovation (KTI) andconducted in cooperation with the Swiss Conference ofWood Industry (HWK Lignum) and the Swiss FederalLaboratories for Materials Testing and Research(EMPA) in Dübendorf.

REFERENCES

1. Hammer, P. (ed.), Acoustic Performance of Medium-RiseTimber Buildings, Dublin, 1998.

2. Cremer, L., and Heckl, M., Körperschall, Springer,Berlin, 1996.

3. Niemz, P., Physik des Holzes und der Holzwerkstoffe,DRW-Verlag, Leinfelden-Echterdingen, 1993.

4. Herrmann, H. J., Hovi, J.-P., and Luding, S. (eds.),Physics of dry granular media, Kluwer, Dordrecht, 1998.

5. Maia, N. M. M., and Silva, J. M. M., Theoretical andExperimental Modal Analysis, Research Studies Press,Taunton, 1997.

FIGURE 2. Summary of standardised impact sound levelsL’n,w obtained for different floors.

FIGURE 3. Operational Deflection Shape of the walkingsurface at a resonance (425 Hz). The bold lines mark thelocations of stiffening ribs at rest.

IIC Performance of Timber Floor Sample with VariousAcoustic Underlay

S. Giglio

PKA Acoustic Consulting, Suite 16, 401 Pacific Hwy, Sydney, 2064, Australia, (sebastiang@pka.com.au)

Australia has seen an emerging trend in recent years of high-rise apartment buildings, including apartments at the upper end ofthe market. Many owners do not wish to install carpet – either for health, aesthetic or cultural reasons. However, hard floorfinishes can give rise to intolerable floor impact noise for the neighbour below. Australia has a short history of high-densityliving so there are no building regulations dealing with floor impact noise, nor are there established solutions for lightweightacoustic floating floors. An investigation was carried out using a 1.44m2 plywood “raft” supported on over 30 differentresilient underlay systems. Impact noise from 50 Hz to 3.15 kHz was recorded although IIC is used as the main descriptor. Itwas found that the IIC performance of a sample was not correlated with the resonant frequency of the floating raft as a mass-spring system, although the best performing system did have the lowest spring resonance. Resonance of the plywood raft itselfappears to play a part in determining performance, suggesting that damping of the raft must be considered as part of any finalsolution. Lightweight fibrous cavity infill was found to be essential for maximising performance.

TEST PROCEDURE

Tests were carried out in a multi-storey apartmentbuilding that was nearing completion, but stillunfurnished. The building was concrete-framed, with170mm thick conventional concrete floor slabs,suspended plasterboard ceilings (~200mm cavity) butno ceiling insulation. The room below the floor beingtested was approximately 100m3. Tapping machinetests were carried out on the bare concrete floor, andthen a large number of tests carried out on the plywoodraft, each with a different resilient underlay system. The plywood raft used for the tests consisted of twolayers of 15mm thick structural grade plywood,1.2x1.2m, glued and screw-fixed together. Thestandard ISO Tapping Machine was used for all thetests. There has been considerable public debate overthe years regarding the Tapping Machine method, withIIC (Impact Insulation Class) as the rating system. IICis acknowledged to be a poor arbiter of floor impactperformance in some cases. It has been used in thisinvestigation for comparison rating purposes in theabsence of a more rigorous system. On the whole, itappeared to provide reasonable ranking, based onsubjective viewing of the spectra.

UNDERLAY MATERIALS

A variety of materials were investigated, including:� Continuous sheets of shredded rubber material,� Glasswool blanket, various densities and thickness� Rockwool blanket,� Needled polyester blanket,� 100mm x 600mm strips of shredded rubber, with

various orientations/arrangements,� Small discrete pads of shredded rubber material,

� Vibration-isolation machine-mounts. Many of the possible solutions that were tried createvoids under the floor, so some of the tests includedlightweight insulation to fill the voids (nominally50mm thick polyester fibre insulation, 10 kg/m3).

DISCUSSION

Some of the findings of the tests are described here.One of the features of the tests on the lightweighttimber floor sample was the very high noise level inthe source room, over 100 dB(A) in some of the tests,and nearly always with a peak at 200 Hz 1/3-octaveband. It appeared that this was a characteristicresonance of the floor sample employed. Airborneflanking had to be accounted for in some test results. The basic floor-ceiling on its own achieved IIC 45.A similar concrete slab, albeit at another location,achieved IIC 34. Based on the relative spectra of thesetwo tests, it is estimated that the suspended ceiling,which had no cavity infill, contributed approximately5 dB reduction of impact noise at low and middlefrequencies compared to a bare slab on its own,increasing up to 12 dB benefit at high frequencies. When the plywood raft was laid on the floor withoutany underlay, IIC 53 was measured. Even though thisis an unrealistic installation method, it suggests thatany separation at all is useful for reducing impactnoise transfer, and that modest levels of impact noiseperformance may be able to be achieved withrelatively thin low-cost materials, provided there is asuspended ceiling installed. This has implications forthe market because it is likely that building regulatorsin Australia will set a relatively low mandatoryminimum performance level, in the range IIC 50-55.On the other hand it is the author’s experience that

strong complaints can be registered even whenmeasured performance on site is over IIC 60. One of the goals of this investigation was todetermine possible means of achieving greater thanIIC 65 with the lightweight timber raft. The goal wasachieved for the prototype floor sample but at theexpense of some complexity of installation.

RESULTS

Table 1 summarises constructions used to achievegiven performance levels for the prototype floor. One of the interesting findings was that all of themineral wool samples of 25mm thickness (50 kg/m3

glasswool, 130 kg/m3 glasswool and 60 kg/m3

rockwool) performed almost within 1 dB of each otheracross the whole frequency spectrum. The resilient underlays that performed the best werethe ones with the lowest mass-spring resonancefrequency. However, systems with identical springfrequency still performed differently. For example, itwas found that changing the orientation andarrangement of the shredded rubber strips could alterthe results considerably, particularly at frequencieswith peak impact sound levels. This appeared to berelated to changes in effective damping of the raft. The machine-mount vibration-isolators offered thebest test results, Figure 1. Varying the mass-springnatural frequency by reducing the number of mountsused under the raft or using single-deflection typesinstead of double-deflection units had very little effecton the resulting impact noise spectra.

Table 1. Summary of constructions to achieve performance.

IIC Resilient underlay systems used to achieveperformance

65+

1. 30mm vibration-isolation mounts loadedto 10-20 Hz, with cavity infill insulation.

2. 50x50x35mm discrete pads of shreddedrubber (550 or 750 kg/m3), loaded to~40 Hz, with cavity infill insulation.

60-64

1. Discrete pads of shredded rubber (550 or750 kg/m3), loaded to ~40 Hz, with orwithout cavity infill insulation.

2. Vibration-isolation mounts loaded to 10-20 Hz, with no other insulation.

3. Strips of shredded rubber, in differentarrangements, with or without cavity infill.

4. 30-50mm thickness of mineral woolinsulation (50 or 130 kg/m3 glasswool, or60 kg/m3 rockwool)

58-61

1. 20-30mm mineral wool insulation (50 or130 kg/m3 glasswool, 60 kg/m3 rockwool).

2. Continuous sheet of 10 or 15mm shreddedrubber (~550 kg/m3).

3. 10 or 20mm thick sheets of needle-punchpolyester fibres (~140 kg/m3).

10203040506070

50 80 125

200

315

500

800

1250

2000

3150

1/3 Octave Band (Hz)

Impa

ct S

ound

Lev

el (d

B)

Bare concrete Best underlay

FIGURE 1. Comparison of tapping machine performancefor bare concrete slab + ceiling and the best vibration-isolation system for the timber raft.

10203040506070

50 80 125

200

315

500

800

1250

2000

3150

1/3 Octave Band (Hz)

Impa

ct S

ound

Lev

el (d

B)

Shredded rubber pads With infill

FIGURE 2. Shows how adding cavity infill can increasenoise at some frequencies (in this case 125 Hz).

The effect of void infill insulation was always toincrease the IIC of the floor system, by 1-6 IIC.However, the infill would usually cause a degradationat one or more frequencies, Figure 2. It ishypothesised that this is due to damped resonancesbeing at a lower frequency than when undamped.

CONCLUSION

It has been possible to achieve IIC 67 for alightweight plywood raft with proprietary isolators, aconcrete slab and a suspended ceiling. This is stillwell short of IIC 80 that can be achieved with carpetand underlay but may be sufficient in many cases.

ACKNOWLEDGMENTS

The author wishes to acknowledge the assistance andinput of his colleagues at PKA in carrying out thisinvestigation. Also, several manufacturers donatedmaterials to enable this investigation to be carried out.

A Case Study on the Estimation of Sound AbsorptionProperties of Worship Buildings

A.Erdem Aknesila, N.Yüğrük Akdağa

aBuilding Physics Department, Yıldız Technical University, 80750, İstanbul, Turkey

In this paper, the studies in the CAHRISMA research project, aiming to conserve the acoustical and visual heritage of worshipspaces in virtual environment is presented from the determination of the acoustical properties of inner surfaces point of view.Three Sinan’s mosques and three Byzantine churches have been selected for this research. Absorption coefficients of innersurface materials in the ancient buildings had to be determinate so as to calculate the acoustical parameters and to be able totransfer the data from the mosques and churches to the software simulation programs. Acoustical properties of the surfaces havebeen determined by estimating their absorption coefficient using bibliographical data.

INTRODUCTION

In order to determinate the acoustical properties ofa room, one of the best way is making measurement. Ifthere is no possibility to determine these data bymeasurements, the only way of evaluation is makingcalculations. It is necessary to obtain the absorptioncoefficients of the materials in this process. Differenttechniques can be used to find out the absorptioncoefficients. Nowadays, data on absorption coefficientsof the usual materials is generally provided by themanufacturers or can be taken from literature.However, to determine the acoustical properties of anancient building, different approaches are generallyrequired, as there is no possibility of determining thesound absorption coefficients for all the materials inthese kind of buildings by measurements andcalculations. Therefore, estimation approach is used. Inthis paper, aiming to emphasise the difficulties on thedetermination of the inner surface absorptioncoefficients of the historical buildings, CAHRISMAProject has been sampled.

The basic goals of the CAHRISMA project are theidentification, revival and conservation of Sinan'sMosques and Byzantine Churches from visual andacoustical heritages point of view in a vitualenvironment. The project incorporates seven teamsfrom six different countries and ten workpackages. Theobjectives of Workpackage 2 which is under theresponsibility of Yıldız Technical University is toselect the worship spaces to work on to collect thenecessary data for constituting the basic database of theproject. Workpackage 2 includes four parts related withfour deliverables. Deliverable 7 covers the architectural

information as well as the data on acoustical propertiesof materials and visual environments of the buildings[1,2].

Determinations of inner surface materials’absorption coefficients which is one of the basic stepsaiming to obtain the data for the investigation is inWorkpackage 2. In this paper, the studies on theabsorption coefficient have been presented.

ACOUSTICAL PROPERTIES OF THEINNER SURFACE MATERIALS

In CAHRISMA Project, because of the mainreasons explained below, it is needed to obtain theabsorption coefficient of inner surfaces materials.� To support the measurements in situ with the

calculations which were done by simulationprograms,

� To realise the audio-visual simulation of thebuildings in software.

By reasons explained below, except few materials,estimation methods have been used to obtain theacoustical properties of the inner surface materials.� Most of the materials of these buildings are ancient

materials. Therefore, there is no possibility to reachthese kind of materials samples for measurements.

� In spite of the fact that, the possibility of finding thesamples of materials, it is difficult to take intoaccount the deformation of materials from theabsorption coefficient point of view in themeasurement procedure.

� The difficulties on the determination of buildingpartitions’ sections.

� Impossibility of getting the ancient materialsamples to make measurements in the laboratory.

After a comprehensive observation on the materials,inside the buildings, the identifications,specifications, section properties of the materialswere elaborated. By using the information obtainedabout materials, absorption coefficient of sixfrequencies have been estimated by matching them

with literature values [3]. Building materials havingdifferent absorption coefficient are represented bydifferent colours on plans and sections co-ordinatedwith coloured lists of absorption coefficients. Inorder to sample the studies, the section of SokulluMosque were presented in Figure 1.

Figure 1. The section of Sokullu Mosque in grey scale and estimated sound absorption coefficients.

CONCLUSION ACKNOWLEDGMENTS

In the simulation programs, inner surfaceabsorption coefficients that were estimated andsampled in this paper were used. Some differencesbetween the simulations and measurements have beenobserved. Especially, at the low frequencies thereverberation times are longer than in situ. On theother hand, in the high frequency range, the simulationis in accordance with the measured reverberationtimes. It can be thought that, the main reason for thesedifferences might be not to have sufficient knowledgeabout the section details of the building partitions.Moreover, the deformation of the materials causedtrouble in obtaining accurate values of the absorptioncoefficient materials from the literature.

The outcome of this study emphasised thenecessity for new advanced measurement techniqueson the surface materials absorption coefficients in situ.

CAHRISMA Research Project (Project No ICA3-1999-10007) is being supported by the EuropeanCommission within the “Confirming of InternationalRole of Community Research INCO-MED” specificprogramme of the Fifth Framework.

REFERENCES

1. Karabiber, Z., “A New Approach to an Ancient Subject:CAHRISMA Project”, 7th International Congress onSound and Vibration, Garmish, Germany, 2000,pp.1661-1668.

2. Karabiber, Z., “A Research on Sinan’s Mosques withinthe 5th Framework Programme of the EuropeanEnvironmental Commission: The CAHRISMA Project”,Tasarım 102, 74-83 (2000).

3. Naylor, G., Rindel, J.H., “Odeon Room AcousticsProgram, Version 2.5, User Manual”, Lyngby, Denmark,1994, pp.91-96.

Method of evaluating of the influence of occurrence vibrations generated by heavy traffic noise on acoustic

conditions in residential buildings

M. Niemas

Department of Acoustic, Building Research Institute, 02-656 Warsaw, 21 Ksawerów St.

The paper presents measurement-calculation procedure, which may be use for estimation of increase in sound pressure level in dwellings radiated from vibrating partitions. The procedure is based on simultaneously twin channel measurements of vibration velocity on partitions and sound pressure level in dwellings. This method using properties of coherence function between has measuring signals. The calculations of the sound radiated from vibrating walls induced by vibro-acoustic sources (e.g. heavy traffic), which has been made by new measure-computational procedure directly from two channel measurements are presented as well.

THE ESSENCE OF THE PROBLEM

The phenomenon of the occurrence of noise in resi-dential premises caused by vibrations on the surfaces of building partitions confining the premises (walls, ceiling, floor) is very complex (Fig. 1). [1]

FIGURE 1 Diagram of the propagation of traffic-induced vibrations to a residential building [1]

The final stage of the phenomenon of transmitting the vibrations energy to the building, that is: emission of acoustic energy by induced to vibrations building partitions confining the residential premises -E p is most important, as it includes the transformation of vibration energy into acoustic energy emitted to the premises. Understanding this phenomenon enables us to develop methods for evaluating the simultaneous occurrence of noise and vibrations in relation to the specific character of transportation sources.

The final effect of the occurrence of vibrations on the partitions in premises given above depends on four acoustic phenomena, namely: 1. the dispersion of vibrations on the building’s struc-

ture, -EB-B

2. inducing the building’s structure to vibrations through the interaction of the base and the buil-ding’s foundation, -ET-B

3. propagation of waves in the base on the source ⇒ building path, -ET

4. inducing the base to vibrations by a passing heavy vehicle (tram, train, bus, truck tractor with semi-trailer, truck), that is generation of a vibration wa-ve, -E• The analysis of the above phenomena will not be

discussed in any more detail, as it is not the subject of the work.

THE MEASURMENT-CALCULATION PROCEDURE

In order to carry out the measurements the follow-ing measurement equipment and additional devices were used: û acoustic calibrator B&K type 4230, û free field microphone B&K type 4165, û microphone preamplifier B&K type 2639, û laser head type POLYTEC OFV-302 with laser vi-

bration meter type OFV-3000, û two-channel portable frequency analyzer B&K type

2144 with programming type 7651 for narrow band analysis by FFT technique,

û tripod for laser head, û connecting cables (BNC-BNC type).

Calculation procedures [1] û The value of the increase in acoustic pressure

∆∆ Lp 1÷200 (increase in the sound pressure level for sound transmission by material path by the exami-ned partitions)

<∆−−=∆

>∆−∆=∆

⋅=∆

ibgdpppbgdppp

bgdppppp

pip

LLLLLL

LLLLL

dBLL

iiiii

iiiii

ii

,,

,

2 ;,γ

(1)

û The sound pressure level for airborne sound trans-mission Lair 1÷200 (using the value of the coherence function)

<=

>=

⋅−=

iii

iiii

ii

bgdpairibgdpair

bgdpairairair

piair

LLLL

LLLL

dBLL

,,

,

2

;

;

;,)1( γ

(2)

û The values: the real sound pressure level Lp 1/3 okt., increase in the sound pressure level

( )

dBpp

pnfpL

ip

kt

kboktk

L

i

k

f

fip

,10

;/log20

0

05.0

0.3/1

⋅=

⋅∑ ∆⋅=

(3)

û ∆∆ Lp 1/3 okt (increase in the sound pressure level for sound transmission by material path by the exami-ned partitions)

( )

dBpp

pnfpL

ip

kt

kboctk

L

i

k

f

fip

,10

;/log20

0

05.0

0.3/1

⋅=

⋅∑ ∆⋅=∆

∆

(4)

û The sound pressure level for airborne sound trans-mission Lair 1/3 okt.,

( )

dBpp

pnfpL

iair

kt

kboktk

L

i

k

f

fiair

,10

;/log20

0

05.0

0.3/1

⋅=

⋅∑ ∆⋅=

(5)

û The total real sound pressure level Lp (1÷160),

( ) dBLk

L

poctkp

,10log1023

1

1.0

1601

.3/1

∑=

=

⋅

÷ (6)

û The total increase in the sound pressure level ∆∆ Lp (1÷160) (total increase in the sound pressure level for sound transmission by material path by the examined partitions)

( )

dB

L

k

L

k

L

p

octkair

octkp

,10log10

10log10

23

1

1.0

23

1

1.0

1601

..3/1

.3/1

∑−

∑=∆

=

⋅

=

⋅

÷

(7)

û The total sound pressure level for airborne sound transmission Lair (1÷160) from the frequency range 1÷160 Hz

( ) dBLk

L

airoctkair

,10log1023

1

1.0

1601.3/1

∑=

=

⋅

÷ (8)

where i2γ , the value of the coherence function for i

frequency band;ipL ,value of the real sound pressure

level for i frequency band; ibgdpL , , the value of the level

of acoustic background in the studied room for i fre-quency band,

kbf , bottom border frequency of k 1/3

octave band, Hz; kt

f , top border frequency of k 1/3

octave band;i

p , absolute value of acoustic pressure in

the next constituent band of k 1/3 octave band; k

n , the

number of measurement bands in the range of the analyzed 1/3 octave band; 0

p , reference value for

acoustic pressure; ∆f, resolution of narrow band analysis (in our case ∆f =1 Hz); i, next frequency band in k 1/3 octave band; k,-next 1/3 octave band

SUMMARY AND CONCLUSIONS

The aim of the work was to demonstrate the possibility of developing and evaluation a method for estimation the influence of vibrations generated by heavy traffic on the acoustic conditions in residential buildings located near communication routes.

The set out goal of the work was achieved, as it was confirmed by measurement that vibrations occurring on confining partitions do influence the value of the sound pressure level in the studied premises.

A measurement procedure was developed based on the statistical method of the coherence function, which made it possible to directly on the spot of the measurement determine the influence of occurring vibrations on the existing sound pressure level in the analyzed premises.

ACKNOWLEDGMENTS

The paper presents the tests and obtained results in the framework of the author’s doctor’s theses, pro-moted by Prof. Jerzy SADOWSKI, Head of Acoustics Department in ITB. Finishing the works was additio-nally financed by KBN in the framework of the pro-moter’s grant no. 7 TO7B 01311

REFERENCES 1. Niemas M.: Ocena wplywu jednoczesnego wystepowania hala-

su i drgan od ruchu komunikacyjnego na warunki akustyczne w budynkach mieszkalnych. Rozprawa doktorska. Promotor Prof. Jerzy Sadowski. Warszawa 1999 (in Polish)

On the use of perforated absorbers in officelandscapes

K. Hagberg, J. Brunskog

Division of Engineering Acoustics, Lund University, P.O. Box 118, SE-223 63 Lund, Sweden

Office landscapes do not act as Sabine-rooms, and therefore special room acoustic criteria have to be formulated and fulfilledin such cases. One such criterion can be that a barrier should have a reasonable insertion and/or transmission loss. Absorbersmade of perforated facings with an air backing is a commonly used product in room acoustic design. This type of product isless effective as an absorber than porous absorber products. Due to the reflection in the ceiling it is usually recommended touse absorbers with alpha close to unity in office landscapes, i.e. porous absorbers. However, in the present pilot studyexperiments indicates that the insertion loss of a barrier is almost the same, whether a perforated or a porous absorber isused. The paper also contains a simple theoretical investigation of the situation.

INTRODUCTION

New office buildings are often built as officelandscapes, or open-plan offices. The acoustics ofthese rooms is often a problem; two colleagues withneighbouring workplaces are likely to disturb eachother when e.g. talking in telephone. Officelandscapes do not perform as Sabine-rooms, andtherefore reverberation time is not a good measurefor these types of rooms. One other criterion that isused instead is that a barrier should have areasonable insertion and/or transmission loss. Absorbers made of perforated facings with an airbacking are a commonly used product in roomacoustic design. This type of product is lesseffective as an absorber than porous absorberproducts. Due to the reflection in the ceiling it isusually recommended to use absorbers with alphaclose to unity in office landscapes, i.e. porousabsorbers. The aim of the present paper is toinvestigate if this rather strict recommendation isnecessary.

MEASUREMENTS

In order to document the attenuation between twotypical workplace locations when different types ofabsorbers were applied to the ceiling, somemeasurements were made in situ in a new officebuilding in Malmö, Sweden. Three types ofabsorbers were tested and, for each case,measurements were carried out both with andwithout a sound-absorbing screen between theworkplaces. The ceilings were applied with the followingtypes of absorbers ; 1) Perforated gypsum (+ 50mm mineral wool behind), 2) Mineral wool 40 mm(soft), 3) Mineral wool 40 mm (stiff). Theabsorption coefficients α for the products areaccording to table 1.

Table 1. Absorption coefficient, α, for thedifferent absorbersFreq. Hz 125 250 500 1000 2000 4000 Case 1 0,55 0,65 0,70 0,65 0,70 0,75 Case 2 0,40 0,85 1 1 1 1 Case 3 0,40 0,65 0,85 0,90 0,95 1

6 m

Screen Loudspeaker Microphone positions

M2 M1 S

6 m

FIGURE 1. Measurement arrangement and ceiling area

The measurement arrangement is described infigure 1. The shaded area in the figure is the area ofthe ceiling included in the measurements, and thisarea is fully applied with the three types ofabsorbers. S is the microphone position in the“sending area” while M1 and M2 are two differentmicrophone positions in the “receiving area”. Thedistance between the loudspeaker and S is 1,2 m.The distance between the loudspeaker and thescreen is 1,4 m. The distance between the screen toM1 is 1,4 m, and to M2 is 2,5 m. The centre of theloudspeaker and the microphones are mounted1,2 m above the floor (corresponding to a personsitting). The mounting height of the ceiling is2,45 m. The screen used was sound absorbing,approximately 30 mm thick and filled with stonewool. It comprised three units each with the width0,8 m and the height 1,8 m. The measurementswere carried out in third octaves within 100 –

5000 Hz. After that, linear sound level differencewas evaluated. The measurement room/landscape was limited inits extension and was not covered with any floorcovering. To minimise reflecting sound from theroom boundaries these were partly covered withmineral wool, see figure 2.

Glass-facade

Open area

11 m 5,5 m

4,3 m

Measurement area acc to fig

1

Mineral wool 100 mm gypsum

FIGURE 2. Boundaries in the measurement room /office landscape

CALCULATIONS

The theory is taken from Thomasson [1], and is amirror source method in an energy sense. A‘sketch’ of the theory is given below. Without screen, the intensity in the receiverposition is

)1(2

20

0 MrrII += (1)

where M describes the mirror sources

10,1,2

2),(0

22

2

≤≤−==

+= ∑

∞

=

qqHr

mqqM

n

m

αβ

βββ

and r is the distance to the receiver, I0 a normalisingintensity, H is the height of the room. An angledependency can be included in q

( ) ( ) .arctan,1β

θθαθ nq n =−=

If the screen is included, we have instead ofequation (1)

( )

102

2

20

0

10 SD

u

ss

s

pp

Mrr

II

−==Π

+Π=

where DS is the screen insulation, calculated fromthe four different paths round the screen (the sourceand receiver is assumed to be located at the floor)

rk

R

ii

i jijjis

∆=+

=Π

ΠΠ=Π ∑∑= =

χχπ

,21

41

,

2

4

1

4

1

where Rij is a interference correction, k is the wavenumber and ∆r is the extra travel path. Finally, the level difference is

( ) ( )02

20 1log10log10 MM

rrD steo +−

+Π= (2)

RESULTS AND DISCUSSION

The sound level differences (linear), see table 2,were received for the different cases whenmeasuring according to figure 1 and 2. Themeasured values are denoted D’ and are comparedwith the theoretical sound level differences Dteo(also linear), equation (2). No angel dependency ofthe absorption or any scattering effects of theceiling were included in the theory.

Table 2. Attenuation (level difference) between thedifferent positions according to figure 1Case 1 2 3scree

nno yes no yes no yes

Rec.pos.

1 2 1 2 1 2 1 2 1 2 1 2

D’(dB)

5 7 12 13 4 6 11 11 5 7 10 12

Dteo(dB)

8 9 12 13 7 10 13 15 7 10 12 14

There are no great differences in the leveldifference for the different cases of absorbers;approximately the cases are equally good. Thus, themeasurement indicates that there are no needs for astrict recommendation of mineral wool absorbers inoffice landscapes. The theory and the measurements have areasonable agreement. However, the theoreticalmodel somewhat overestimates the influence of thescreen for the mineral wool absorbers (case 2 and3) but not for the perforated absorber. Assumingthat the theory is correct in overall sense, thisoverestimation can be due to angel dependency ofthe absorption or scattering effects of the ceiling, asthese phenomena were not included in the model.

REFERENCES

1 Thomasson, S-I. Beräkning av insättningsisolering förskärmar inomhus, Report TRITA-TAK-8101, TekniskAkustik, Kungl. Tekniska Högskolan, Stockholm, 1981(in Swedish)

Investigation of Sound Environment in Historic SitesS.-pin Huanga and R.-ping Laib

aDepartment of Architectural Engineering, Kao-Yuan Institute of Technology, No.1821,Chung- Shan Rd ,Luju, Kaohsiung county, Taiwan, R.O.C.

bDepartment of Architecture, National Cheng-Kung University,No.1, Ta-Hseuh Road, Tainan city, Taiwan, R.O.C.

The purpose of this research is about sound environment in historic sites. Sound environment includes outside soundenvironment and inside sound environment. We discuss time history of outside sound environment, types of noise and specialsoundscapes by field measurements. Besides we measure some inside acoustical environment of historic buildings by usingprevious instruments. Then we scientifically analyze some acoustical properties such as reverberation times and sound pressurelevel distribution by field measurements.

INTRODUCTIONHistoric site is a part of cultural heritage. Peoplealways ignore the environment in or around historicsites is also a part of cultural heritage. Theenvironments in or around historic site create leisurelyactivities such as walking exercise, sightseeing,ceremony and business. Sound environment ischanging by above different activities. Soundenvironment is also different in any era.The study investigates sound environments in historicsite in Tainan City. Tainan City is the origin of thedevelopment of Formosa. Tainan is recognized as theoldest city in Taiwan. Tainan city has many historicsites such as Confucian Temple, Chih-kan Tower,Erkunshen Cannon Fort, Remains of Taiwan City, TaTien Hou Temple and Chi Tien Wu Temple. Weinvestigate sound environment in Confucian Templeand Erkunshen Cannon Fort, Confucian Temple is themost important heritage in Tainan. The temple wasoriginally built in 1665. And Erkunshen Cannon Fortmeans “Eternal Castle”, it was built in 1876.

MEASUREMENTSOutside Sound Environment

Confucian Temple locates at the center of Tainan City.The study investigates outside sound from 7:00 to17:00. And measuring Leq per continued five minutes.We choose usual day and holiday to compare the soundenvironment. The instruments are two sound pressurelevel meters and a digital recorder. The results showenvironmental sound pressure levels at court are from50dB(A) to 79dB(A) in Figure 1. The average is about60 dB(A). Figure 2 shows sound pressure levels arefrom 58dB(A) to 85dB(A). The average is about 65dB(A) in front garden more 5dB(A) than at court . Thenoise sources are from morning exercises, school’sannounces, aircraft, people, tourists and cars. Somesoundscapes are from birds, Chinese traditional music,

drums, fountain and selling bell. Erkunshen CannonFort locates at a new urban area and close to An-pingharbour. This site is extremely quiet now. But I think itwill be much noisy because the area is developing.Figure 3 show environmental sound is low. Soundpressure levels are from 42 dB(A) to 61 dB(A). Theaverage is about 48 dB(A) . Sound inside entrance archis louder in Figure 4, because it is a passing tunnel.Sound reflect make echo sound. Sound environment inErkunshen Cannon Fort is much lower than inConfucian Temple. Sound environments of two sitesare obviously different.

Sound Distributions at CourtWe are interested in sound propagation at main court inConfucian Temple. Confucian memorial ceremony isheld at this court on September 28th every year. So wemeasure sound pressure levels at 15 positions at halfcourt. Then we compare those values to discuss sounddistribution. Figure 5 shows the positions andloudspeaker. And Figure 5 also shows the differentvalues at 500Hz. The court is an open space. Sosound propagation reduces quickly about 16dB(A) to18dB(A) at 500Hz. It means when memorialceremony is held that the sound pressure levels couldbe not enough for audience. The phenomenon is moreserious at 1000Hz to 4000Hz. It also influences theperformance of outside concert or live.

Inside Acoustical Measurement inConfucian Temple

The ceiling of Chinese traditional architecture iscomplex and beautiful. It is constructed by woodenframe and roof tile. Sound tracing is more various.We first research room acoustics in Chinesearchitecture. Reverberation times are measured byusing sound analyzer. The results are shown in Table1. We use two sound sources, clap and balloon. Itseems good for room acoustics.

CONCLUSIONS

1.Sound investigation could help us to realize activitiesin historic sites.

2.Sound recording would become a part of heritage infuture.

3.historic sites are threat by traffic noise, aircraft noiseand visitors’ noise. Some soundscapes need torecorded and reserved.

4.Sound distributions at open court are not good fortraditional music, ceremony, live or speech.

5.Reverberation times in Confucian Temple are

measured and the results show good.6.Sound environment in Erkunshen Cannon Fort is

quiet at daytime and nighttime.7.Sound environment in historic sites are interesting

and deeply influence people’s life.

Table 1. Reverberation times in temple (Unit: second)125Hz 250Hz 500Hz 1000Hz 2000Hz 4000Hz

Clap 2.45 1.98 2.09 2.01 1.96 2.04Balloon 2.67 2.34 2.23 1.83 1.88 1.87

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FIGURE 3. Environmental sound in ErkunshenCannon Fort

FIGURE 2. Environmental sound at front garden inConfucian temple

FIGURE 1. Environmental sound at court inConfucian temple

FIGURE 4. Environmental sound inside entrancearch in Erkunshen Cannon Fort

Investigation of Noise Control Methods for Simulated Rain

S. Giglio

PKA Acoustic Consulting, Suite 16, 401 Pacific Hwy, Sydney, 2064, Australia, (sebastiang@pka.com.au)

Traditionally, profiled metal sheeting is used as a cost-effective construction method for many types of buildings in Australia.This has an inherent disadvantage in that, when untreated, it can give rise to considerable levels of rain noise in occupiedspaces. One manufacturer of roof insulation and plasterboard has set out to systematically test various solutions. A stand-alone single-room building was purposely constructed in an open space and sprinklers installed to simulate natural rain.Although the contact velocity of the water drops was much lower than for natural rain, the rain noise spectra obtained indoorswas very close to that of natural rain. On this basis, insertion loss testing was carried out for various roof-ceiling acoustictreatments.

TEST RIG

A single-room building was constructed in an opencar park behind the manufacturer’s factory. Theexternal walls are double-skin and use separatedtimber stud frames. The roof framing is supported bythe outer wall only. The room volume is 50m3 withouta ceiling and 40-42m3 with a ceiling installed. To dateonly low pitch (3-5˚) metals roofs have been tested.Plan dimensions of the room are 4.5 x 3.5m. The floorslopes in two directions so ceiling height varies from2.5-2.75m. Entry to the room is via back-to-back solidtimber doors with acoustic seals. There are four sprinklers installed at a height of 1.6mabove the roof. These have a square spray pattern with75˚ included angle and are fed from a 6,000 litreholding tank via a 4 kW pump. Intensity of simulatedrainfall is determined by collecting the runoff from theroof and measuring the time taken to fill a 75 litre bin. The test procedure involves measuring soundpressure levels within the room, which can then beconverted to sound power level. Acoustic treatment ofthe roof then allows the insertion loss (synonymouswith noise reduction in this case) to be determined.

RAINFALL INTENSITY

A common design criterion used for non-tropicalareas of Australia is 10mm/hour rainfall intensity. Thetest rig operates at approximately 400mm/hr.However, recorded dB(A) sound levels are 14 dB(A)lower than expected for natural rain of this intensity,Dubout [1]. This is explained by the relatively lowdrop height (1.6m) compared to that for natural rain.McLoughlin et al [2] establish a relationship betweenwater drop height and rain sound level and thiscorresponds very well with the discrepancy noted here.

FIGURE 1. The roof profile used as the base case hasheight 29mm, pitch 190.5mm, and metal thickness 0.42mm.

SPECTRA OF RAIN NOISE

4045505560657075

125 250 500 1k 2k 4k 8k1/1 Octave Band Centre Frequency (Hz)

Soun

d Le

vel (

dB)

Dubout McLoughlin CSR

FIGURE 2. Rain spectra standardised to 75 dB(A).

Figure 2 shows the simulated rain noise spectraobtained from the test rig, compared to natural rain(Dubout) and McLoughlin et al (4.43m drop heightwith a 5.5mm diameter drop). The comparison spectrahave been standardised at 75 dB(A). The roof profileused by Dubout was similar to the base case used here,except the height was 41mm and the metal 0.8mmthick. The profile used by McLoughlin et al was0.9mm thick with a 38mm deep trapezoidal profile.

ROOF INSULATION

The most common application of insulation to ametal roof is in the form of a fibrous glasswool blanket50-55mm thick, with foil bonded to one side. This isusually installed for thermal reasons, with acoustics asecondary consideration. It is evident, however, thatthis technique also provides significant damping of theradiated acoustic energy from the roof. In most areasof Australia the foil is installed facing down. Intropical areas of Australia and Asia the foil is installedface up, against the metal roof. The most common foil

grade is 250 gsm (grams/m2). Figure 3 summarises theresults with roof insulation. It can be seen that:

� Even the least substantial insulation was still ableto provide 11 dB(A) reduction of rainfall noise.

� The law of diminishing returns appears to be atwork in this instance, providing a maximum of19 dB(A) noise reduction that can be achieved byroof insulation alone, that is, without a ceiling.

520 600840

880 9001320

300

02468

101214161820

25 40 50 55 60 75 100 120Thickness of insulation (mm)

Inse

rtion

loss

(dBA

)

Rockwool Glasswool Polyester

FIGURE 3. Insertion loss of each insulation material, testedwith foil face down. The weight in gsm of the sample isshown on the figure. The Rockwool samples were 2000 gsmand 1800 gsm, respectively.

This analysis has shown that thicker and/or morebulky insulation provides no additional noise reductionbeyond a certain limit. Further, such materials canresult in an undesirable uneven external appearance ofthe roof cladding. It was determined that with the foil face installedagainst the metal roof, rather than the fibrous material,the insertion loss decreased by 2 dB(A) when testedwith the 55mm 600 gsm glasswool sample. A test was also carried out with lightweight foil(150 gsm compared to 250 gsm) bonded to 75mm 720gsm glasswool. This showed negligible performancedifference.

RAIN NOISE REDUCTION PROVIDEDBY A CEILING

Tests were also carried out to determine the insertionloss when a ceiling is installed below the roof. Testarrangements included metal-rod suspension systemwith butterfly-clips, resilient ceiling hangers, directly-fixed ceiling, set plasterboard ceiling, plasterboard

ceiling tiles in an exposed metal grid and mineral fibretiles in an exposed grid. Findings can be summarisedas follows:

� A 13mm set plasterboard ceiling with a metal-rodsuspension achieved insertion loss of 22 dB(A).This is 3 dB(A) better than the best roof insulationon its own and 7 dB(A) better than the typical roofinsulation used.

� The insertion loss of the ceiling can be added tothat of the roof insulation when a metal-rodceiling suspension is used. This is probably dueto the butterfly clip, which appears to filter outsufficient high-frequency vibration.

� Using rubber-element ceiling-hangers provided asmall additional benefit with regards to rain noise,compared to the metal-rod suspension system.

� When a ceiling is installed, then insulation laidover the ceiling is nearly as effective as roofinsulation, when comparative insulation materialsare used. Acoustic absorption in the ceiling spaceis obviously an important factor in the insertionloss, not only roof damping.

� Doubling the plasterboard ceiling lining resultedin a further 3-4 dB(A) noise reduction.

� When both ceiling insulation and roof insulationare installed, significantly higher noise reductionis achieved than using either on its own, althoughnot quite equal to the sum of the insertion losses.

� Plasterboard ceiling tiles in an exposed gridprovided insertion loss only 3 dB(A) worse thanthe set plasterboard ceiling, but mineral fibreceiling tiles were 7 dB(A) worse.

� The insertion loss of the direct-fixed plasterboardwas 7 dB(A) worse than the suspended ceiling.

ACKNOWLEDGMENTS

The author wishes to thank David Carne, of CSRResearch, for providing encouragement and assistance,as well as permission to publish the results of theinvestigation. Thanks also go to CSR Gyprock andCSR Bradford companies for being prepared toactively push forward building industry research inAustralia, particularly with regards to acoustic issues.

REFERENCES

1. Dubout P., Journal of Sound and Vibration, 10(1) 144-150 (1969).

2. McLoughlin J., Saunders D.J., Ford R.D., AppliedAcoustics, 42, 239-255 (1994).

The Contribution of Thermal Insulation Constructive Solutions to the Improvement of

Sound Insulation of FloorsL. Bragançaa, S. Silvaa, J. Patríciob

a) - UM, Dep. Eng. Civil, Azurém, 4800 Guimarães, Portugal; e-mail: braganca@civil.uminho.pt; sms@civil.uminho.ptb) - LNEC, Av. do Brasil, 1700 Lisboa, Portugal; e-mail: jpatricio@lnec.pt

In Portugal, the accomplishment of thermal insulation in buildings with national regulations is a pre-requisite for the licensing ofhousing buildings. This fact makes necessary the use of adequate products in external walls, ceilings and floors. In case of floors, the use of thermal insulation layers may significantly contribute to the improvement of sound insulation againstairborne and impact noise, yielding constructive solutions that better allow buildings to accomplish national noise regulations. Generally, in what concerns airborne noise, the use of floors with enough mass per unit area or significant internal loss factors is aprimary condition for the accomplishment of regulations. But regarding impact noise, that increase in mass is not always enoughbeing resilient floor coverings then needed. And what about the contribution of insulation thermal products usage? What is gained interms of noise insulation? Could this usage be defined as a noise insulation safety margin both in terms of airborne and impact noise? In this paper, some results and constructive recommendations are given based on the experience of using insulation thermal layersin floors, enhancing thatl these aspects may contribute to the improvement of noise insulation of floors regarding impact and airbornenoise as an important complement to floor coverings efficacy.

INTRODUCTIONThe noise insulation between apartments is becoming

more important as typical housing noise sources,essentially televisions and HiFi equipments, are gettingmore powerful. Nowadays, the use of floor softcoverings with resilient properties, like cork or carpets, isgetting out of fashion for health reasons. So, thedevelopment of new constructive solutions becomes thennecessary to reduce sound propagation between spacesincreasing, consequently, the respective sound insulation.

In Portugal, to accomplish the existing buildingsthermal legislation [1], it is necessary to use thermalinsulation materials in both external and internal buildingenvelope, namely in walls, ceilings and floors.

The use of thermal insulation products in floors whichseparate spaces with different environments can lead to asignificant energy saving, especially when one of theapartments is acclimatized and the other is not (or evenin the case of very different indoor conditions).

The use of a thermal insulation layer on the floorsbetween apartments can also contribute to a betteracoustical performance of these partitions constructiveelements enabling them with good acousticalcharacteristics concerning insulation to both airborne andimpact sound.

FLOOR NOISE INSULATIONIn order to accomplish the limits imposed for floors by

Portuguese noise regulations [2] concerning the airbornesound reduction index, R’w, and the impact sound index,L’n,w, it is necessary to prevent the occurrence of soundtransmission through all the structural elements of thebuilding [3].

Table 1 lists the noise insulation requirements forfloors separating apartments and for the partitionsbetween a commercial area and an apartment.

Table 1 – Portuguese legal requirements for floors [2]

Noise insulationbetween:

R’w (dB) L’n,w (dB/oit)

Apartments � 48 � 70Apartments andcommercial areas � 55 � 55

Table 2 lists the noise insulation indices obtainedexperimentally in several residential blocks, studied bythe authors. These values are typical for non-insulatedfloors.

Table 2 – Typical Portuguese floors noise insulation

Floor Type R’w (dB) L’n,w (dB/oit)Floor 1 53 82Floor 2 47 90Floor 3 48 88Floor 4 49 89Floor 5 52 87Floor 6 51 86

Comparing the measured typical noise insulationindices for floors with those established in the legislation,it can be seen that neither the impact noise nor theairborne noise indices fit the Portuguese legalrequirements. So, solutions to improve noise insulationare needed. This is usually achieved using resilient layersplaced between the structural slab and the floor covering.In this traditional approach the resilient layer has onlyacoustical properties. Figure 1 schematically shows thistype of solution.

3

5

4

21

Figure 1 – Traditional solution for the impact noise reduction

Thus, a thermal insulation material can also be used asa resilient layer, associating its thermal properties with itsacoustical characteristics. There are several materialsappropriate for the purpose: cork aggregate, glass ormineral fiber, expanded and extruded polystyrene, andexpanded polyurethane. From thermal insulation point ofview, a 2 to 3 cm thickness layer is enough to guaranteeits efficiency. The extruded and expanded polystyrene iswidely used as thermal insulation as it has an excellentcost/efficiency ratio. Figure 2 shows how to use thismaterial in terms of thermal and acoustic aspects. Apartthe material, the main difference between this procedureand the one outlined in Figure 1 is the addition of aplastic film to avoid any possible contacts between theconcrete in the base of the covering, and the structuralslab.

6

12

5

43

Figure 2 – Thermal insulation material used to improveacoustic behaviour of a slab

GLOBAL SYSTEM EFFICIENCYIn Table 3, it is presented a comparison between the

results for typical Portuguese floors (Table 1) and thoseobtained with experimental measurements performed onsimilar floors having thermal insulation products appliedin the way Figure 2 illustrates.Table 3 – Noise insulation of floors with thermal insulation

Floor Type R’w (dB) L’n,w (dB/oit)Table 1 Fig. 2 Sol. Table 1 Fig. 2 Sol.

Similar to floor 1 53 55 82 67Rehab. of floor 2 47 52 90 70Similar to floor 3 48 52 88 68Similar to floor 4 49 51 89 69Rehab. of floor 5 52 57 87 68Rehab. of floor 6 51 53 86 67

The improvement obtained with this constructive solutiondepends on some factors and can be estimated between 15dB to 20 dB for the impact sound and between 2 dB to 5 dBfor airborne sound.

As it can be verified in Table 3, the proposed solution isadapted for floors separating apartments. For thoseseparating apartments and commercial and service areas thelegal requirements are not fitted with this solution. For thistype of floors it is necessary to adopt, additionally, asuspended ceiling, with a minimum height of 15 cm,partially filled with an absorbent material, as for example,mineral fiber. This solution generally yields an additionalairborne sound insulation varying between 3 dB to 6 dB.

CONCLUSIONSSeveral experiments performed “in situ” and in

laboratory confirm that the adoption of thermal insulationconstructive solutions can contribute to the improvementof sound insulation of floors and to an adequate acousticcomfort in dwellings. The described recommendationslead to sound reduction indices that meet the Portugueselegal requirements for floors. The proposed solutions aresimple and economical and, therefore, valid for the mostcommon horizontal partitions between dwellings.

REFERENCES

1. Regulations of thermal characteristics in buildings, Decree-Law nº 40/90,February, 6 th

2. Portuguese noise regulations – Decree-Law n.º 292/00, November 14th

3. Silva, M. P. – Acoustics in Buildings, LNEC, ITE 8, Lisbon 19954. PATRÍCIO, J. V. – Acoustic performance of non-homogeneous

floors regarding impact sound in buildings: simulation model “Ph.D. Thesis”). LNEC, Lisbon, 1999

- Floor Covering - Base for the Covering - Resilient Layer

- Structural Element

- Ceiling Finishing

- Floor Covering - Base for the Covering - Plastic Film - Resilient Layer

- Structural Element

- Ceiling Finishing

An Experimental Study Of Devices For The Control OfReverberation By Using Textile Material

J. Ramis, J. Alba, J. Redondo and J.M. Bravo

Departamento de Física Aplicada, Escuela Politécnica Superior de Gandia, 46730 Grao de Gandia, Spain

In this work results from measurements of absorption coefficient for different textile and porous materials are presented. Theporous sheets should be applied for the control of reverberation in the range of high frequencies. Most porous sound absorptivesheets do not present attractive surfaces . The solution more used is to cover the absorbent material with a textile material withacoustically transparent but pleasing facing. We analyse the influence of superficial density of textile and the density of poroussheets in the air cushion (plenum). The results are interesting from the point of view of architect that it is interested in roomboundaries with the same appearance but with different acoustical properties.

INTRODUCTION

The acoustic conditioning of auditoriums must bedevelop solutions in accordance with overall estheticalconcept. Most porous sound absorptive sheets do notpresent attractive surfaces. The solution more used isto cover the absorbent material with a textile materialwith acoustically transparent but pleasing facing.Nevertheless, the absorption coefficient of thecomplete device change. In this work, we present someexperimental results related with the influence ofcharacteristics of different fabric. The results areinteresting from the point of view of architect that it isinterested in room boundaries witch the sameappearance but which different acoustical properties.Figure 1 show the typical configuration of the device.

fabricair

absorbent

wall

FIGURE 1. Typical configuration

DETERMINATION OF THEABSORPTION COEFFICIENT IN A

REVERBERATION ROOM

The procedure for the determination of theabsorption coefficient in a reverberation room isdescribed in [2]. It can be obtained from measurementsof the reverberation time with and without materialsample. Limitations and uncertainly related with thismethod are discussed in [1] and [3].

MEASUREMENTS

We present some results of a experimental study forthe characterisation of devices made with five kind offabric with the same appearance but different physicalproprieties. The absorbent material used in this workare rockwool sheets of different densities. Devices with an air cushion (plenum) ranging from2.5 to 15 cm, totally and partially filled of rockwool of40, 70 y 90 Kg/m3 ser measured. With the plenumpartially filled, the rockwool as been placed next to thewall and next to the fabric. Next, we show someresults. For this presentation, we have defined an arbitrarygraduation, from 1 to 5, related with the degree ofacoustical transparency (1 for the more transparent and5 for the less) . In figure 2 we compare the absorptioncoefficient of the five fabric mounted on an air cushionof 5 cm (without rockwool). In figure 3 is shown thesame configuration but with the plenum totally filledof rockwool of 90 kg/m3 of density. The effect of therockwool of 40 kg/m3 is shown in figure 4. In figure 5is presented the result for the fabric the moretransparent fabric and in figure 6 for the lesstransparent.

0 ,0 0

0 ,1 0

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0 ,3 0

0 ,4 0

0 ,5 0

0 ,6 0

0 ,7 0

0 ,8 0

0 ,9 0

1 0 0 1 0 0 0 1 0 0 0 0f (Hz )

α

1 2

3 4

5

FIGURE 2. Absorption coefficient for five fabric with anair cushion of 5 cm without material absorbent

0 ,0 0

0 ,1 0

0 ,2 0

0 ,3 0

0 ,4 0

0 ,5 0

0 ,6 0

0 ,7 0

0 ,8 0

0 ,9 0

1 ,0 0

1 0 0 1 0 0 0 1 0 0 0 0f (H z )

α

12345

FIGURE 3. The same as in figure 2, but with the air cushiontotally filled with Rockwool of 90 kg/m3.

0 , 0 0

0 , 1 0

0 , 2 0

0 , 3 0

0 , 4 0

0 , 5 0

0 , 6 0

0 , 7 0

0 , 8 0

0 , 9 0

1 , 0 0

1 0 0 1 0 0 0 1 0 0 0 0f ( H z )

α

12345

FIGURE 4. The same as in figure 3, but with the air cushiontotally filled with Rockwool of 40 kg/m3

0,00

0,20

0,40

0,60

0,80

1,00

100 1000 10000

α

Air cushion

Rockwool-90 kg/m3

Rockwool-40 Kg/m3

FIGURE 5. Absorption coefficient of fabric number 1 withrockwool of 90 Kg/m3 and 40 Kg m3.

0,000,200,400,600,801,00

100 1000 10000f (Hz)

α

Air cushion

Rockwool-90kg/m3

Rockwool-40kg/m3

FIGURE 6. The same as fig. 5, but for the fabric number 5.

CONCLUSIONS

If the permeability of fabric decrease, the absorptioncoefficient is incremented. The introduction ofabsorbent material in the air cushion (plenum),produces an increment in the absorption of the system.,and all devices behaves similarly. When the aircushion is partially filled with absorbent material, it ismore efficient if it placed next to the plate. The morepermeable is the fabric the more is the increase inabsorption coefficient after introduction of theabsorbent material.

REFERENCES

1. E. Cremer et Al., Principles and Applications of RoomAcoustics, London: Applied Science Publishers (1982).

2. UNE 74-041 (and ISO 354). Medida del coeficiente deabsorción en cámara reverberante.

3. 3.Alba, J. et AL, Incertidumbre en la técnica de medida de laabsorción en cámara reverberante.. Tecniacústica 99. Ávila(Spain).