NATu !«,S' OF S-A«40 i TECH R 1 C

ill iilliilll iiilillll nil III! Ill NiST

PUBLICATIONS

NisrUnited States Department of Commerce

Technology Administration

National Institute of Standards and Technology

NIST Special Publication 864

Technical Digest

Symposium on Optical Fiber

Measurements, 1994

Sponsored by the National Institute of Standards and Technology

in cooperation with the IEEE Lasers and Electro-Optics Society

and the Optical Society of America
NISTTechnical Publications

Periodical

Journal of Research of the National Institute of Standards and Technology -Reports NIST

research and development in those disciplines of the physical and engineering sciences in which the Institute is active. These include physics, chemistry, engineering, mathematics, and

computer sciences. Papers cover a broad range of subjects, with major emphasis on measurement methodology and the basic technology underlying standardization. Also included

from time to time are survey articles on topics closely related to the Institute's technical and

scientific programs. Issued six times a year.

Nonperiodicals

Monographs -Major contributions to the technical literature on various subjects related to the

Institute's scientific and technical activities. Handbooks -Recommended codes of engineering and industrial practice (including safety

codes) developed in cooperation with interested industries, professional organizations, and

regulatory bodies.

Special Publications -Include proceedings of conferences sponsored by NIST, NIST annual

reports, and other special publications appropriate to this grouping such as wall charts, pocket

cards, and bibliographies.

Applied Mathematics Series -Mathematical tables, manuals, and studies of special interest to physicists, engineers, chemists, biologists, mathematicians, computer programmers, and others

engaged in scientific and technical work.

National Standard Reference Data Series -Provides quantitative data on the physical and

chemical properties of materials, compiled from the world's literature and critically evaluated.

Developed under a worldwide program coordinated by NIST under the authority of the National

Standard Data Act (Public Law 90-396). NOTE: The Journal of Physical and Chemical

Reference Data (JPCRD) is published bi-monthly for NIST by the American Chemical Society

(ACS) and the American Institute of Physics (AlP). Subscriptions, reprints, and supplements are

available from ACS, 1155 Sixteenth St., NW, Washington, DC 20056.

Building Science Series -Disseminates technical information developed at the Institute on

building materials, components, systems, and whole structures. The series presents research

results, test methods, and performance criteria related to the structural and environmental

functions and the durability and safety characteristics of building elements and systems.

Technical Notes- Studies or reports which are complete in themselves but restrictive in their treatment of a subject. Analogous to monographs but not so comprehensive in scope or

definitive in treatment of the subject area. Often serve as a vehicle for final reports of work

performed at NIST under the sponsorship of other government agencies.

Voluntary Product Standards -Developed under procedures published by the Department of Commerce in Part 10, Title 15, of the Code of Federal Regulations. The standards establish

nationally recognized requirements for products, and provide all concerned interests with a

basis for common understanding of the characteristics of the products. NIST administers this program in support of the efforts of private-sector standardizing organizations.

Consumer Information Series -Practical information, based on NIST research and experience,

covering areas of interest to the consumer. Easily understandable language and illustrations provide useful background knowledge for shopping in today's technological marketplace.

Order the above NIST publications from: Superintendent of Documents, Government Printing

Office, Washington, DC 20402.

Order the following NIST publications -FIPS and NISTIRs-from the National Technical

Information Sen/ice, Springfield, VA 22161.

Federal Information Processing Standards Publications (FIPS PUB) - Publications in this series

collectively constitute the Federal Information Processing Standards Register. The Register serves as the official source of information in the Federal Government regarding standards

issued by NIST pursuant to the Federal Property and Administrative Services Act of 1949 as amended, Public Law 89-306 (79 Stat. 1127), and as implemented by Executive Order 11717

(38 FR 12315, dated May 11, 1973) and Part 6 of Title 15 CFR (Code of Federal Regulations).

NIST Interagency Reports (NISTIR)-A special series of interim or final reports on work

performed by NIST for outside sponsors (both government and non-government). In general,

initial distribution is handled by the sponsor; public distribution is by the National Technical

Information Service, Springfield, VA 22161, in paper copy or microfiche form.
NIST Special Publication 864

Technical Digest—Symposium on Optical

Fiber Measurements, 1994

Digest of a symposium sponsored by the

National Institute of Standards and Technology

in cooperation with the IEEE Lasers and Electro-Optics Society

and the Optical Society of America

September 13-15, 1994

National Institute of Standards and Technology

Boulder, Colorado 80303-3328

Edited by

G.W. Day

D.L. Franzen

R.K. Hickernell

September 1994

U.S. DEPARTMENT OF COMMERCE, Ronald H. Brown, Secretary

TECHNOLOGY ADMINISTRATION, Mary L. Good, Under Secretary for Technology

NATIONAL INSTITUTE OF STANDARDS AND TECHNOLOGY, Arati Prabhakar, Director
National Institute of Standards and Technology Special Publication 839

Natl. Inst. Stand. Technol., Spec. Publ. 864, 225 pages (Sept. 1994)

C0DEN:NSPUE2

U.S. GOVERNMENT PRINTING OFFICE

WASHINGTON: 1994

For sale by the Superintendent of Documents, U.S. Governnnent Printing Office, Washington, DC 20402-9325
PREFACE

This Digest contains the manuscripts of 50 papers, 10 invited and 40 contributed, presented at the

eighth Symposium on Optical Fiber Measurements, September 13-15, 1994, in Boulder, Colorado.

The invited papers were selected from about 50 nominations. The contributed papers were selected

from 57 submitted.

The most significant theme of the Symposium is the present importance of polarization

measurements. More than 20 percent of the papers are concerned either with polarization

measurements, especially polarization mode dispersion and polarization dependent loss, or with the

characterization of fibers and components with special polarization properties. Optical time-domain

reflectometry (OTDR) measurements are another important theme, as they have been since the first Symposium in 1980, and the characterization of optical fiber amplifiers continues to be important.

Nonlinear processes in fiber seem to be a growing interest.

Approximately two thirds of the papers come from outside the United States, up from about half at the seventh Symposium. Ten countries are represented in the program.

G.W. Day

D.L. Franzen

R.K. Hickernell

Boulder, Colorado

September 1994

Except where attributed to NIST authors, the content of individual sections of this volume has not been reviewed or edited by the National Institute of Standards and Technology. NIST therefore accepts no responsibility for comments or recommendations therein. The mention of trade names in this volume is in no sense an endorsement or recommendation of the National Institute of Standards and Technology.

iii
SYMPOSIUM COMMITTEE

D.L. Franzen, NIST, General Chair

G.W. Day, NIST Program Chair

N. Channon, York

P. Di Vita, CSELT

R.K. Hickernell, NIST

M. Jinno, NTT

W.T. Kane, Corning

A.D. Kersey, NRL

R.B. Kummer, AT&T Bell Laboratories

P.S. Lovely, Photon Kinetics

S.A. Newton, Hewlett-Packard

S. Pollitt, NPL

S.B. Poole, University of Sydney

W.A. Reed, AT&T Bell Laboratories

P.R. Reitz, AMP

C. Saravanos, Siecor Canadian Cable

H.H. Yuce, Bellcore

iv
CONTENTS

Noise figure measurements in optical fibre amplifiers (invited)

S.B. Poole, Univ. Sydney 1

Optical amplifier noise figure determination by signal RIN subtraction

F.W. Willems, J.C. van der Plaats, AT&T Network Systems Nederland;

C. Hentschel, E. Leckel, Hewlett-Packard GmbH 7

The intricacies of an EDFA round-robin: First observations from COST-241 and

RACE "COMFORT

"

F.W. Willems, AT&T Network Systems Nederland 11

Crosstalk in 1.3 ^m praseodymium fluoride fibre amplifiers

S.C. Fleming, Univ. Sydney 15

Measurement and compensation of gain peaking in multi-amplifier

links (invited)

E.L. Goldstein, Bellcore 19

Out-of-band on-line monitoring of the optical network (invited) ' E. Cottino, AET; M. Artiglia, CSELT 25

Multi-100 Gbit/s optical waveform measurement using sum-frequency-generation

optical sampling with subpicosecond supercontinuum pulses

H. Takara, S. Kawanishi, T. Morioka, K. Mori, M. Saruwatari, NTT Transmission

Systems Laboratories 31

Characterizing modal interference in field installable single-mode fiber connectors

incorporating short fiber stubs

D. O. Harris, R.A. Throckmorton, Siecor 35

Performance limits of coherent OTDR due to optical nonlinear effects (invited)

H. Izumita, Y. Koyamada, S. Furukawa, I. Sankawa, NTT Telecommunication

Field Systems R&D Center 39

Windowing technique for accurate measurement of low reflectances by OTDR

L. Ducos, LCIE; P. Facq, IRCOM 45

Noise error in OTDR splice loss measurement

F. Caviglia, P. Ricaldone, CSELT 49

Effects of fiber parameter mismatch on uni-directional OTDR splice loss measurement

J. Warder, M.J. Li, P. Townley-Smith, C. Saravanos,

Siecor Canadian Cable Plant 53

Practical improved splice loss estimation for single-mode optical fibers

H. Tanji, S. Ito, International Digital Communications 57

V
OTDR Trace Analysis for fibre characteristics evaluation: A "bi-directional" approach

M. Artiglia, CSELT; A. Chiantore, G. Ferri, FOS; A. Rossaro, CSELT 61

Determination of the inhomogeneity of a single-mode fiber along its length by

bidirectional OTDR measurements at two wavelengths

C. linger, W. Stocklein, Siemens 65

Comparison of algorithms for evaluating fibre attenuation longitudinal uniformity

A. Chiantore, G. Ferri, FOS; M. Artiglia, P. Di Vita, CSELT 69

1 m spatial resolution measurement of distributed Brillouin frequency shift in

single-mode fibers

T. Horiguchi, T. Kurashima, Y. Koyamada, NTT Telecommunication

Field Systems R&D Center 73

The development of an optical fibre attenuation standard

N.R. Haigh, BICC; T.C.E. Jones, NPL 77

Microwave based measurements of optical fibers

R.M. Atkins, AT&T Bell Laboratories 81

Characterization of in-fiber Bragg gratings (invited)

K.O. Hill, F. Bilodeau, B. Malo, S. TheriauU, D.C. Johnson, J. Albert,

T. Kitagawa, Communications Research Centre 85

Characterization of optical components using high resolution optical reflectometry

techniques (invited)

W.V. Sorin, D.M. Baney, S.A. Newton, Hewlett-Packard Laboratories 91

Polarization-maintaining fiber alignment using tunable polarization-maintaining

fiber Fabry-Perot interferometers

Y. Bao, K. Hsu, CM. Miller, Micron Optics 97

Tapered fiber diameter measurements

P. Nebout, N. Godbout, S. Lacroix, X. Daxhelet, J. Bures, Ecole

Polytechnique de Montreal 101

Characterization of silicon optical bench waveguides (invited)

C.H. Henry, H.H. Yaffe, AT&T Bell Laboratories 105

Diagnosis of silica-based waveguides by using a jaggedness-free OLCR

K. Takada, H. Yamada, M. Horiguchi, NTT Opto-Electronics Laboratories .... 109

Measurements of loss and mirror reflectivity in semiconductor optical waveguides:

A European interlaboratory comparison experiment

E. Gini, ETHZ; L.H. Spiekman, H. van Brug, TU Delft; J.-F. Vinchant,

Alcatel Alsthom Recherche; S. Morasca, F. Pozzi, C. De Bernardi, CSELT;

R. Zengerle, W. Weiershausen, W. Noell, DBP Telekom; L. Thevenaz,

A. Kung EPFL; A. Enard, N. Vodjdani, Thompson CSF 113

vi
White-light spectral analysis of lithium niobate waveguides

C.F. McConaghy, K.F. Hugenberg, D. Sweider, M. Lowry, Lawrence Livermore

National Lab; R.A. Becker, Integrated Optical Circuit Consultants 117

Frequency response measurements of integrated-optic electrodes

K.F. Hugenberg, P.D. Sargis, C.F. McConaghy, Lawrence Livermore

National Lab 121

Fiber diameter from the backscattered light pattern

J. Baines, NPL: C. Frehill, Dublin City Univ.; K. Raine, NPl 125

Fiber geometry: Results of an international interlaboratory measurement comparison

T.J. Drapela, D.L. Franzen, M. Young, NIST 129

Contact and non-contact interferometry for optical fiber diameter measurement

and calibration

C. Brehm, O. Mercereau, Alcatel Alsthom Recherche; G. Lipinski,

G.P. Vailleau, LNE Laboratoire National d'Essais 133

Light scattering fiber/coating eccentricity measurements for thin coatings

D. H. Smithgall, AT&T Engineering Research Center 137

Transverse optical scattering measurements of the core radius of high-A

fibers with 1-nm resolution

P.N. Saeta, B.I. Greene, D.J. DiGiovanni, A.E. Miller,

AT&T Bell Laboratories 141

Measurement of fiber core position for fiber array

K. Ozawa, S. Ogawa, H. Ishida, Y. Hattori, Sumitomo Electric Industries 145

Polarization mode dispersion: Definitions, measurements and statistics (invited)

N. Gisin, Univ. Geneva 149

Mode coupling model for the Jones matrix eigenanalysis method

D.A. Nolan, D.Q. Chowdhury, Corning 155

Study on a conservation quantity in PMD measurements

O. Aso, I. Ohshima, H. Ogoshi, Furukawa Electric Co., Ltd 159

Quantitative comparison between the Jones matrix eigenanalysis & the wavelength

scan method for polarization mode dispersion measurement: A simulation study

D.Q. Chowdhury, D.A. Nolan, Corning 163

Polarisation mode dispersion measurement using Fourier analysis: Investigation

of mode coupling behaviour in fibers

A.J. Barlow, EG&G Fiber Optics 167

vii
Inter-laboratory polarization mode dispersion measurement study

W.B. Gardner, AT&T Bell Laboratories; A.J. Barlow, EG&G; K.A. Emig, Corning;

S. Grindstaff, AT&T Fitel; P. Hernday, Hewlett-Packard; W.C. Jones, Bellcore;

T. Olson, Alcatel; D.K. Roland, SpecTran 171

Characterization of dispersion compensating fibers (invited)

A. M. Vengsarkar, A.E. Miller, AT&T Bell Laboratories 175

Tailoring of dispersion compensation fibers with high compensation ratios up to 30

R. Bonefi, TU Dresden; J. Vobian, FTZ Darmstadt; S. Unger,

J. Kirchof, IPHT Jena 181

Characterization of single polarization fiber

M.J. Messerly, R.E. Budewitz, B.K. Nelson, 3M Fiber Optics Laboratory;

R.C. Mikkelson, Macalester College 185

New direct measurement technique of polarization dependent loss with high

resolution and repeatability

N. Mekada, A. AL-Hamdan, T. Murakami, M. Miyoshi, Santec 189

The statistics of polarization dependent losses

N. Gisin, Univ. of Geneva 193

Effects of fiber nonlinearity and polarization mode dispersion on

long-distance transmission systems with optical in-line amplifiers (invited)

T. Imai, M. Murakami, T. Takahashi, NTT Transmission Systems Laboratories . . 197

Highly repeatable measurement of nonlinear refractive index by cross phase ^ modulation method using depolarized pump light

T. Kato, Y. Suetsugu, M. Takagi, E. Sasaoka, M. Nishimura, Sumitomo

Electric Industries 203

Automated fiber nonlinear coefficient measurements using self-phase modulation method

Y. Namihira, KDD R&D Laboratories; Y. Tanaka, M. Mure, Oyokoden Lab. Co. 207

High-accuracy Brillouin gain spectrum measurements of single-mode fibers L. Thevenaz, M. Nikles, P. Robert, EPFL Swiss Federal Institute of Technology . 211

Guided acoustic-wave Brillouin scattering as an ultrasensitive probe of optical

fiber geometry

B. I. Greene, AT&T Bell Laboratories; P.N. Saeta, NIST/JILA 215

AUTHOR INDEX 219

viii
Noise Figure Measurement in Optical Fibre Amplifiers

Simon Poole

Optical Fibre Technology Centre, University of Sydney, NSW 2006, Australia

1 Introduction

Since the first reports of erbium-doped optical fibre amplifiers (EDFAs)^'^ the key measurements

required have been those of amphfier gain, output power and noise. The first two are relatively easily

measured using a tunable source, attenuator and a spectrum analyser, with only minor compensation

for the effects of AmpUfied Spontaneous Emission (ASE) at low input signal powers being required.

Amplifier noise measurements are, however, more difficult due to the interaction of the signal with

the ASE.

Amplifier noise is a critical parameter for in-line and pre-amplifiers. In the former, it is the

cumulative effect of the noise which can limit system performance whilst in the latter additional noise

leads to a decrease in receiver sensitivity. Noise contributions from power ampUfiers are generally

less critical, except for AM-TV distribution systems. Nevertheless, accurate measurements of noise

figure are essential to both amplifier manufacturers and users to ensure optimum system performance.

In this paper, the origins of noise in fibre amplifiers are reviewed and the Noise Figure of an amplifier

defined (Section 3.) Section 4 describes a range of techniques which have been developed for noise

figure measurement in a production/test environment whilst Section 5 looks at the rapidly-developing

area of in-service noise figure monitoring. The relative strengths and weaknesses of the various

techniques are compared in Section 6. Throughout this paper, reference will be made primarily to the

Erbium-doped Fibre Amplifier (EDFA) for operation at around 1.55|am although the techniques

outlined can also be used, with minor modifications to account for the different excited-state lifetime,

for noise measurements in Praseodymium-doped Fluoride Fibre Amplifiers (PDFFAs) around 1.3|Lim.

2 Origin of Noise in Amplifiers

Gain in an EDFA is obtained by optically pumping the electrons in erbium ions from the \^/2 groundstate to the "^IjjQ metastable, followed by stimulated emission when a signal photon is incident on the

excited atom. This inversion can be achieved by pumping directly into the metastable level with light

at 1.48|im^ or indirectly through light at QSOnm" Energy transfer mechanisms involving Er/Yb ions^ can also be used and, for the purposes of the consideration of amplifier noise, can be considered to have the same performance as 980nm pumping.

However, if stimulated emission does not occur, the excited erbium ions will eventually decay

spontaneously to the ground state with the emission of a photon which has random phase and direction

and with a wavelength anywhere within the fluorescence spectrum of the amplifier. The photon may

be captured by the fibre and amplified, leading to amplified spontaneous emission, which is unpolarised and the major contributor to noise in EDFAs.

The amplifier design (co- or counter-propagating pumps, single or double pumps, etc) and the choice

of pump wavelength can significantly affect the amplifier noise figure. Generally, amplifiers pumped

at 980nm have a lower noise figure^ (theoretical minimum =3dB but typically around 3 - 4 dB in practical amplifiers) than those pumped at 1480nm (typically 5-7dB) since the former achieve a higher

inversion of the erbium ion population than the latter.

The noise sources in a receiver for a signal that has been amplified by an optical amplifier can be

summarised as:

1
Signal shot noise due to the intrinsic thermal noise of the source. This is usually small although for accurate definition of the amplifier Noise Figure (see below) this term must be included.

Spontaneous shot noise from the amplifier spontaneous emission. Again, this is usually negligible.

Signal-spontaneous beat noise caused by the mixing of the signal and amplified spontaneous emission

which falls within the detector bandwidth. In a well-designed amplifier, this is the dominant

term. Spontaneous-spontaneous beat noise due to mixing of the amplified spontaneous emission with itself on a detector. This is the dominant noise term at very low signal powers but for a welldesigned amplifier this contribution can generally be ignored.

Signal excess noise caused by multiple reflections from components inside the amplifier. This is primarily a problem for systems which employ narrow line-width sources of the type

commonly found in AM TV distribution systems.

The noise figure (NF) of an amplifier is defined in terms of the signal to noise ratios measured with an ideal optoelectronic receiver as the optical bandwidth (and hence the contribution from spontaneousspontaneous beat noise) is reduced to an infinitesimally small value and is defined

as NF = Urn ^ where Bg is the optical bandwidth of the receiver. For optical measurement

techniques, it is generally necessary to simplify this definition by setting the measurement conditions

such that signal spontaneous beat-noise is the dominant factor, giving a simplified noise figure of

NF " ^^ASE vvhere p^^g is the ASE power density with the same polarisation and wavelength as

(G-1)/7v

the signal, G is the optical gain and hx) is the photon energy. In practice, the receiver optical

bandwidth is finite and the effect of spontaneous-spontaneous beat noise should be included.

However, this contribution is typically less than the signal-spontaneous noise for optical bandwidths

of around 5nm.

The exact definition of amplifier noise figure, including which minor terms need to be included, is

still under discussion in various standards groups worldwide''. However, for the purposes of

comparing the various measurement techniques, the exact definition is of secondary importance.

3 Noise Figure Measurement Techniques

Optical measurement techniques for amplifier noise measurement all require a measurement or estimation of the ASE noise component within the signal band. This measurement can be made

optically through techniques such as spectral subtraction and interpolation (Section 3.1), polarisation

nulling (Section 3.2), and pulse recovery techniques (Section 3.3) or electrically through RIN

subtraction (Section 3.4) A number of points are common to all of the techniques. In particular, the

effects of reflections and connections losses from the measurement or stimulus instrumentation need

to be minimised to ensure accurate, repeatable measurements. In practise this is often achieved by

splicing the amplifier under test to the measurement instrumentation to minimise any reflections^.

3.1 Spectral Interpolation

Spectral interpolation is probably the simplest measurement technique for ASE estimation and

uses an optical spectrum analyser to measure the ASE noise either side of the signal, with the

ASE at the signal wavelength being interpolated either linearly or through a curve using a

priori knowledge of the shape of the noise spectrum around the wavelength of interest (Fig

1). The simplicity of the technique allows measurements to be made rapidly but it is highly

sensitive to source spontaneous emission, particularly in highly-saturated amplifiers.

Furthermore, it cannot be used to characterise amplifiers (eg line or pre-amplifiers) which

contain any form of narrow-band output filter. Spectral hole burning also limits the technique

2
to amplifiers which are homogeneously broadened, although this is not a problem for current generation EDFA and PDITAs.

3.2 Polarisation Nulling

The polarisation-nulhng technique^ uses a linearly polarised input signal and the spectra at the

amplifier output measured for both the polarisation of the transmitted polarised light and for the orthogonal polarisation. Since the ASE is randomly polarised, whilst the input signal is linearly polarised, measure the in-band ASE power on the orthogonal polarisation to that of

the input signal can be measured. The experimental apparatus for this measurement is shown

schematically in Figure 2.

Residual source laser sidebands can introduce significant error into the measurement and care must be taken to ensure maximum suppression of the sidebands using an in-line bandpass

filter'^. It should also be noted that, as for all of the techniques outlined, the source power

should be varied through an external attenuator rather than through a varying drive current. This ensures that the ratio of source power to sideband noise is constant throughout the

measurement.

Polarisation dependent gain or loss and polarisation mode dispersion can introduce errors into

the polarisation nulling technique, particularly for measurements in highly saturated amplifiers.

The first two create a polarisation-dependence of the ASE spectrum, whilst the latter can limit

the degree of nulling of the amplified signal.

3.3 Pulse Recovery

In the pulse recovery technique" the principle of measurement is very simple; The input

signal is chopped at a low frequency (typically 2-300Hz) but with a fast (<1 pSec) fall-time. Immediately after the signal is blocked, the ASE at the output of the amplifier is measured

using a on opto-electronic converter and a digitising oscilloscope. A typical trace is shown in Fig 3. Since the ASE is measured when the signal source is blocked, there is effectively

infinite rejection of the signal and any associated sidebands. Furthermore, the ASE is measured at the signal wavelength, thus obviating the need for any interpolation of measured

values either side of the measurement wavelength. The technique has recently'^ been extended

to provide accurate measurement of noise figure in the saturated regime. Measurements have

been made with up to 15dB of gain compression with an accuracy of +/-0.5dB.

3.4 RIN Subtraction

Electrical measurement techniques use a comparison of the measured of the relative intensity

noise (RIN) of a laser before and after amplification to provide a direct measurement of the

noise contributed to a signal by the optical amphfier. This should be contrasted with optical

technques which use an indirect measurement, with the measured parameters related to the

actual amplifier noise, generally through some approximate expression. RIN is measured on

an electrical spectrum analyser and is the noise measured electrical power per unit bandwidth

normalised to the detected dc electrical power. For amplifier characterisation, it is udsual to choose a region of the laser noise spectrum away from any spectral features and the amplifier

noise can then be determined by simply subtracting the noise of the source as measured

without the amplifier from that measured when the amplifer is present. To obtain accurate measurements, a well characterised, very high quality optical receiver is required and

considerable care must be taken to achieve repeatable results.

3
4 In-service noise-figure measurement

In service measurement of amplifier noise is a relatively recent development and hence measurement

techniques for this are somewhat less developed than for laboratory or production measurements.

Forward propagating ASE at wavelengths away from the signal wavelength using the technique shown

schematically in Figure 4 has been demonstrated". Here a small proportion of the ASE is tapped off

and passed through a narrow band filter before measuring on a photodiode. To prevent cumulative

errors due to ASE amplification in multiple amplifiers in the chain a band reject filter is placed in the

signal line to reject the wavelengths used for monitoring. Despite the approximations required a measurement accuracy of +/- O.OSdB with a reproducibility of +/- 0.03dB was achieved for a wide

range of input signal powers at a carefully controlled wavelength. The technique is extendable to other

wavelengths but at a reduced level of accuracy due to the approximations used.

A second configuration uses a circulator to monitor the backward propagating ASE'"*''^ (Fig 5). This

has the advantage in that the ASE can be measured at the signal wavelength although the technique

as described suffers from the need to assume uniform inversion along the amplifier. Since this is unlikely to be achieved in practice, there is a potentially significant error due to this assumption. The

authors report that noise figures accurate to within IdB of the measured forward propagated noise

figure were achieved for input powers below -6dBm. The system is advantageous in that the measured

power levels are significantly higher than for the previous example but suffers form the need to use

a circulator, a device which is not yet available in commercial quantities or specifications.

A third alternative is the use of a shaped slit/integrating photodiode combination'^ which relies on the

detection of side-scattered spontaneous emission along the fibre length, with the effective capture angle

for the detector weighted as a function of position along the fibre. This technique is shown

schematically in Figure 6. Calculations indicate an absolute accuracy of +/-0.08dB, although this was

difficult to establish experimentally due to scatter in the measured results to which it was being

compared.

5 Discussion

A number of techniques for amplifier noise figure measurement have been developed which can be

broadly broken down into optical domain and electrical (RF) domain measurements. Whilst the latter are able to take into account all aspects of amplifier noise, including internal reflections, and provide

a direct measurement of the amplifier noise figure, the former are more generally used due to their

simplicity, ease of implementation and reduced measurement times. Table 1 below summarises the

relative advantages and disadvantages of the techniques discussed above.

Technique Advantages Disadvantages

Signal Interpolation Quick, easily implemented Lack of traceability

Polarisation Nulling Quick, accurate Polarisation dependent loss/gain introduces

errors

Pulse Recovery Good in saturation More complex technique

RIN Subtraction Direct measurement of amp noise Complex, can be operator dependent

In-service measurements Real-time Cost, complexity, are they necessary?

Table 1 Comparison of amplifier noise figure measurement techniques

There is still some discussion as to the need for in-service measurements of noise figure, since

degradation of the pump laser (the most likely reason for amplifier performance degradation) is most

easily detected by monitoring the amplifier output power, rather than the noise figure.

4
Despite the number of techniques which have been developed, there is no one procedure which has

gained universal acceptance. This is due in part to the lack of direct, objective comparisons between

the different techniques. To the authors' knowledge, only two studies have been reported'^ '^ which

have attempted to directly compare results obtained by both optical and electrical measurement

techniques. This is clearly an area requiring further investigation and where the efforts of standards

working parties can be of immense value to both the amplifier manufacturer and the amplifier users. Further agreement on the definition of amplifier noise figure may also be required before a definitive

set of test methods can be developed.

6 Acknowledgments

The author would like to thank many colleagues world-wide for assistance in the preparation of the manuscript.

Particular thanks go to Steve Newton of Hewlett Packard, Paul Morkel of STC Alcatel and Richard Laming of

the ORC, Southampton University for the supply of background information and to Ralph Betts and Steve

Frisken of Photonic Technologies Pty Ltd for useful discussions.

7 References

1 R.J.Mears, I.M.Jauncey, L.Reekie, D.N.Payne, "High-gain rare-earth doped fiber amplifiers at 1.54 /jm", Proc OFC/IOOC87, paper WI2

2 E.Desurvire, J.R.Simpson, P.C.Becker "High-gain erbium-doped travelling-wave fiber amplifier" Opt.Lett.,

11, 1987, pp.888-890.

3 E.Snitzer, H.Po, F.Hakimi, R.Tuminelli, B.C.McCollum, "Erbium fibre laser amplifier at 1.55pm with pump

at 1.49pm and Y b-sensitised oscillator" Proc OFC'88 paper PD-2.

4 R.LLaming, S.B.Poole, E.J.Tarbox "Pump excited-sate absorption in erbium-doped fibres" Optics Lett., 13,

p.1084, 1988

5 S.G.Grubb et al Proc 0FC'91 paper PD7.

6 R.Olshansky "Noise figure for erbium -doped optical fibre amplifiers", Electron. Lett., 24, 1988, pl363

7 eg WG6 of IEC-T686 on Optical Fibre Amplifiers

8 D.Baney, C.Hentschell, J.Dupre "Optical fiber amplifiers - Measurement of gain and noise figure" 1993 HP

Lightwave Symposium,

9 J.Aspell, J.Federice, B.Nyman, D.Wilson, D.Schenk "Accurate noise figure measurements or erbium-doped

fibre amplifiers in saturation conditions" Proc OFC'92, paper ThA4, 1992

10 T.Kashiwada, M.Shigematsu, M.Nishimura "Accuracy of noise figure measurement for erbium-doped fiber

amplifiers by the optical method" Tech Dig Symposium on Optical Fibre Measurements 1992, pp209-212

11 D.M.Baney, J.Dupre "Pulsed source technique for optical amplifier noise figure measurement" Proc ECOC'92, paper We P2.1 Ipp 509-512

12 K.Bertilsson, P.A.Andrekson, B.E.Olsson "Noise figure of Erbium Doped Fibre Amplifiers in the Saturated

Regime" IEEE Photon. Technol. Lett.,vol 6, pp. 199-201, 1994.

13 H.Masuda, K.Aida, K.Nakagawa "Noise figure monitoring of a cascaded in-line erbium-doped fiber

amplifier" IEEE Photonics Tech. Lett., 5 1993, pp.1436 - 1438. 14 Y.Sato, Y.Yamabayashi, K.Aida "Monitoring the noise figure of EDFAs with circulators via backward

amplified spontaneous em ission " Proc Topical Mtg on Optical Amplifiers and their Applications, Yokohama,

Japan, 1993, paper MD-14

15 Y.Sato, Y.Yamabayashi, K.Aida "Noise figure monitoring of optical amplifiers via backward amplified

spontaneous emission" IEEE Photonics Technol Lett, 6, 1994, pp. 203-204.

16 H.Masuda, K.Aida "A novel noise figure measuring and monitoring method for erbium-doped fiber

amplifiers with high resolution (O.Oldbj and short response time (50pS)" Proc Topical Mtg on Optical

Amplifiers and their Applications, Yokohama, Japan, 1993, paper MD-17

17 S.Nishi, M.Saruwatari "Comparison of noise figure measurement methodsfor erbium-dopedfiber amplifiers" Proc ECOC '93, paper Tu3

18 F.W.Willems et al "Comparison of saturated EDFA NF determination with the ESA and OSA methods" Proc OFMC, Turino, paper 5.2, 1993.

5
Figure 1 Spectrum Analyser trace of amplifier ASE

and signal showing interpolation markers

TLS

PM

HMISSA/ttSXIA

»04.»iwz«iloig _

STATE t*-

COMTPIOtltn 1

OSA »

HFTUSOA nxeo7

^OUMIZATICIN

Figure 2 Schematic of apparatus for polarisation nulling technique

(from [81)

Figure 3 Typical measurement trace for pulse recovery technique (from ref 12)

9PF

POg ffl 0 9ft>iinLD K>FA-2

Figure 4 In-service noise-figure measurement technique

due to Masuda [13]

• Output Pon

Input Pon

Figure 5 Circulator-based in-service noise figure measurement [141

Transmissicin f^**'

Rtsef IsolatDf Coupter

(a)

Signal Ugtrt ASE

Isolator

0.98-(UH LD L. PO

SJit >SE

Transmission

Fltwr

Bandpass

RIter

Figure 6 Schematic of integrating photodiodetechnique for in- service noise figure monitoring [16]

6
Optical Amplifier Noise Figure Determination

by Signal RIN Subtraction

F.W. Willems and J.C. van der Plaats

AT&T Network Systems Nederland B.V.

P.O. Box 18, 1270 AA Huizen, The Netherlands

C. Hentschel and E. Leckel

Hewlett-Packard GmbH

Herrenberger Strasse 130, Boblingen, Germany

Abstract: A novel way to determine the noise 2 Theory

figure of a saturated optical amplifier is pre- sented, based on subtracting the EDFA input

and output Relative Intensity Noise. The re- sulting noise figure is baseband frequency re- solved and includes the contribution of interferometric effects caused by multiple reflections and

Rayleigh scattering in the ampUfier.

The noise terms in a system using saturated optical amplifiers, taking into account all coupling

losses are [3] signal shot noise, relative intensity noise caused by interferometric effects in the

EDFA and signal spontaneous beatnoise, assuming a shot noise limited source (i.e. no signal

laser RIN) and negligible receiver thermal noise.

1 Introduction

The Noise Figure (NF) of an optical amplifier can, in principle, be determined using an Optical

Spectrum Analyzer (OSA) by measuring the DC

Amplified Spontaneous Emission (ASE) power

and applying a theoretical relationship between

this power and the noise figure [1]. Wavelength

dependence is easily obtained with a tunable ex- ternal cavity laser. Noise from Multiple Path

Interference (MPI) at the signal wavelength [2] can be added if the reflectivities in the amplifier are known; however, this requires access to the

EDFA internal components.

Alternatively, an Electrical Spectrum Analyzer (ESA) can be used to measure input and

output Signal-to-Noise Ratios (SNR); subtracting the SNRs leads to the NF by definition. In

this case, noise from MPI is automatically in- cluded. However measurement of the SNR at several electrical baseband frequencies is time

consuming. In this paper we present a measure- ment method that has none of these disadvantages: the RIN subtraction (RINS) method. It combines high measurement speed, full automation potential and complete noise figure characterization.

_ Ahvusp G* -1

Nshot,9ignal = 2eT]iG*7]oLpPoBe (1) — - (rjiG*rf,LpPofBe (2)

Nmpiif) = RINmpi{f){ViG*VoLpPofB, (3)

with Be the electrical noise measurement bandwidth, Ugp the spontaneous emission coefficient

(usp > 1), e the electron charge, h Plancks constant, V the optical signal frequency, rji the input

coupling efficiency of the EDFA, 770 the output

coupling efficiency of the EDFA, L the coupling

efficiency between the EDFA output and the re- ceiver, G* the internal optical gain of the EDFA,

Po the optical power just in front of the EDFA,

p = ^ the receiver responsivity (77 being the

detector quantum efficiency, including the coupling efficiency), RINmpi{f) the relative intensity noise caused by incoherent interferometric

effects in the EDFA.

The noise figure is defined by

NF =

SNRin

sm,out (4)

If we assume by definition that the detector is ideal, i.e. 7/ = 1 and L = 1, and if we define

7
G = T)iG*rio we arrive at

NF{f) = (RINmM) + RINs-sp) 2^ + ^ (5)

The spectrum analyzer measures the total noise

RINtotal 1 including the signal laser diodes noise

RINldH) which has to be subtracted and thus:

NFU) = (RINtotalif) - RINLDif))

^

+ ^

(6)

A detailed error analysis for the OSA method

used in the HP 81600 EDFA test system shows

that the total measurement uncertainty of the

EDFA noise figure is 0.4 dB, which can be re- duced to 0.25 dB if fusion splicing of the EDFA

to the test system is applied [4]. The accu- racy of the RIN subtraction method is limited by

the amplitude accuracy of the modulated power

trace, which is 2 dB electrical.

3 Experiment

The NF of a specially prepared EDFA, see fig- ure 1, has been measured with the OSA and the

RIN-subtraction method respectively. For the

OSA method the HP 81600 Series 200 EDFA

test system was used, incorporating a tunable

external cavity laser (TECL), polarization in- sensitive optical spectrum analyzer and polarization nuUing capabilities. The input power was -2 dBm, leading to a saturated condition in which

ASE power is much smaller than the signal out- put power; that is important because the optical

power measurement during the RIN procedure

does not distinguish between signal power and

ASE and secondly the RIN subtraction method

presumes that spontaneous-spontaneous beatnoise is absent. For the RIN subtraction method

a 1556 nm DEB laser with a linewidth of 1.5 MHz and a RIN below -160 dB/Hz were used.

An HP 71400 lightwave signal analyzer set with

RIN measurement software was employed.

Figure 1 shows the noise figure of the EDFA

as a function of electrical baseband frequency

for two settings of the Variable Optical BackReflector (VOBR): R = -68 dB and R = -15

dB. If J? = —68 dB, the cavity reflectance is small enough to avoid interferometric noise and

EDFA module Variable Erbium Doped Fibre Bacl^eflector

Signal

In '980 nm Pump

Signal Out

Figure 1: EDFA with internal Rayleigh Mirror and

external Variable Optical Backreflector.

the noise figure is more or less independent of the

electrical frequency. The data from the RIN subtraction method for R=-15 dB follow the curve

as predicted in [5], indicating that the cause of

the noise figure increase at lower frequencies is phase to intensity noise conversion due to optical cavities in the EDFA and that the length of

the cavity is smaller than the coherence length

of the laser. In this case the cavity is formed by

the effective Rayleigh Mirror of the EDFA and

an external Variable Optical Backreflector.

A simplified model for the interferometric

noise assuming only incoherent mixing, is given

by

NF RINs—sp Au

eff f2 A+ 1

2hiy G

(7)

With Reff the effective cavity reflectance from,

in this case, amplified Rayleigh scattering and

the discrete mirror, including polarization alignment, Au the FWHM of the source linewidth

(Lorentzian model). With an amplified Rayleigh

scattering of -41 dB, we find, assuming perfect

polarization matching of the direct and doubly

reflected signals that i?e// ~ ~28 dB; using this

value in equation 7 gives the solid curve plotted

in figure 2. The data from the OSA method are on a straight line by definition and approach the

tail of the RIN subtraction data.

4 Conclusions

An EDFA noise figure measurement method has

been described that is able to resolve Interferometric phase to intensity noise conversion due

to the presence of optical cavities in a saturated

EDFA. It is based on the subtraction of input

8
a

O

16

14

7^ 12

10

TT—I—I—I—I—I—I—r—

r

K = -15 tf^iAftA

1= -68 dB

J I J L J I L

100

Frequency [MHz]

Figure 2: Noise figure as a function of electrical baseband frequency using the OSA method (straight

lines) and the RIN subtraction method at a signal in- put power of -2 dBm at 1556 nm; the solid curve was

obtained from modeling the reflections with equation

7.

[2] W. I. Way, C. Lin, C. E. Zah, L. Curtis, R. Spicer, and W. C. Young, "MultipleReflection-Induced intensity noise studies

in a lightwave system for a multichannel AM-VSB television signal distribution,"

IEEE Photonics Technology Letters, vol. 2,

pp. 360-362, May 1990.

[3] N. A. Olsson, "Lightwave systems with optical amplifiers," IEEE Journal of Lightwave

Technology, vol. 7, pp. 1071-1082, July 1989.

[4] D. Baney, C. Hentschel, and J. Dupre, "Optical fiber amplifiers - measurement of gain

and noise figure," Lightwave symposium,

200 Hewlett-Packard, 1993.

[5] J. C. van der Plaats and F. W. Willems,

"Relative intensity noise increase caused by

amplified signal redirected Rayleigh Scattering in Erbium doped fibers," in Digest

of Conference on Optical Fiber Communications '94, San Jose, pp. 158-160, Feb. 1994.

Paper WM6.

and output Relative Intensity Noise that can be

measured with a calibrated spectrum analyzer

and thus gives at once the NF frequency re- solved. The method gives results that are comparable with the OSA method when the influ- ence of MPI is negligible.

Acknowledgement

This work was partly supported by the Commission of the European Union under the RACE II program, project R2062, "Components for Fibre

to the Home (COMFORT)".

References

[1] J. Aspell, J. F. Federici, B. M. Nyman, D. L. Wilson, and D. S. Shenk, "Accurate noise fig- ure measurements of erbium doped fiber amplifiers in saturation conditions," in Digest

of Conference on Optical Fiber Communications '92, San Jose, pp. 189-190, Feb. 1992.

paper ThA4.

9

The intricacies of an EDFA Round-Robin:

First observations from COST-241

and RACE "COMFORT"

F.W. WiUems

AT&T Network Systems Nederland B.V.

P.O. Box 18, 1270 AA Huizen, The Netherlands

Abstract: Fundamental differences between the

ESA and OSA noise figure measurement method and

the more "trivial" problems of measuring an EDFA's

NF, such as calibration and definitions are discussed,

based on observations made during a "mini" roundrobin along five European laboratories. The differ- ences between results can be attributed to calibration

of the optical resolution bandwidth and absolute optical power as far as OSA is concerned. Major problems with the ESA method are in the proper handling

of source noise levels.

1 Introduction

COST 241 "Characterization of Advanced Fibres for the new Photonic Networks" working group 3, study

group 2 is currently performing an EDFA round-robin

among the participants of this group. The RACE R2062 "COMFORT", "Components

for fibre to the home" project is developing EDFAs

for analog applications. One task is to identify the

measurement procedure for noise behavior taking into account all effects that can influence the system operation.

The common task of both COST 241-WG3-SG2

and COMFORT is to make a well defined measure- ment protocol. This requires a good knowledge of all the problems that are encountered in doing the measurement and those are often not foreseeable. Therefore we choose as a first approach to accompany the

round-robin EDFA (RREDFA) with a person that

could monitor how the measurements were performed

and where the difficulties arose. We will report first observations and point out the pitfalls we discovered

so far in the round-robin.

2 Noise Figure measurement

Methods: OSA and ESA

In terms of input and output signal-to-noise ratio, the

noise figure of an optical amplifier as mea.sured by an

electrical spectrum analyzer, is defined by

The signal-to-noise ratio at the EDFA input is deter- mined by three noise sources: signal laser RIN, re- ceiver shot noise and thermal noise; to make a comparison with the optical measurements possible, we

require shot noise limited operation of the system,

i.e. low intrinsic RIN of the laser and high input

powers in the photodetector.

It can be shown [1] that the ASE level is related

to equation 1 by

assuming that the signal spontaneous beatnoise dominates (i.e. we use an optical filter in front of the

system's detector or we operate the EDFA in saturation).

3 Fundamental differences between OSA and ESA

It has been pointed out previously [2] that differ- ences in noise figure measurements by the ESA and

OSA method are often caused by different definitions. However by identifying the proper relation- ships between the parameters that are measured by

both methods, one can usually end up with comparable noise figures (see section 2). Unfortunately two

fundamental differences in the ESA and OSA method

have been identified. The first is related to the fact that it is impossible with the OSA method to measure

the noise at the optical carrier frequency, whereas this

is exactly what one does with the ESA method. The

second is even more intricate and plays a role in satu- rated amplifiers only; it is related to the fact that the

photon statistics of a saturated, non-linear EDFA dif- fers from the statistics in the linear regime, (i) Inter- ferometric noise caused by incoherent multi path in- terference is always present in EDFAs even when the

utmost care is taken to prevent discrete reflections.

Its source is Rayleigh scattering in the Erbium Doped

Fibre itself. This kind of noise can only be detected

by the ESA method and can influence the noise fig- ure significantly [3]. (ii) The equations that relate the

ASE power level to a signal-to-noise ration degradation have been derived in the regime where the EDFA

11
is not saturated [1]. It is show that when the real pho- ton statistics in a saturated EDFA is taken into ac- count by calculating the photon master equation, the

noise figure can be significantly higher than the one

calculated from the DC ASE power level [4] [5]. Differ- ences in noise figures from optical measurements that

are converted to the CNR degradations by the con- ventional rules (see section 2) are mostly attributed

to measurement errors, but could well be caused by

phenomenon just described.

Both fundamental observations imply that the optical measurement method can not, from first principles, give a correct noise figure in all cases.

4 Measurement results

Since for the ESA method, knowledge of the exact gain is not required, some labs did not implement

the gain measurement at this time. Therefore we will restrict ourselves in this paper to presentation of the

noise figure measurement results. To compare the ESA and OSA values, we made use of the conversion

equations mentioned in section 2 realizing that these

are in principle wrong according to the observations

in section 3. The device used for the round-robin was a double 1480 nm pumped, temperature and pump current

stabilized EDFA with input and output isolators to ensure that operation of the device is not dependent on the residual reflection in the measurement system.

Further requirements are long term stability and in- dependence of temperature. We measured the noise

figure of the device over a period of six months and

the values stayed in a 0.2 dB range. In doing the measurement we took care to use each time the same cal- ibration procedure. The noise figure of the RREDFA

is also hardly dependent on temperature as can be

seen in figure 1. First we will show the influence of two successive procedure changes we implemented. Refer to

figure 2. Curve A was our first measurement us- ing an optical spectrum analyzer with polarization

nulling [6]. The ASE power was measured in an optical resolution bandwidth of 0RBW=1 nm according

to the OSA. When this bandwidth was calibrated us- ing a tunable external cavity laser over a wavelength

range it turned out that the "real" bandwidth var- ied between 0.965 at 1520 nm to 0.92 nm at 1560 nm. At a resolution bandwidth of 0.5 nm the situ- ation was much better, since the variation was now

between 0.505 and 0.485 nm which has hardly any

influence on the noise figure anymore: for this particular spectrum analyzer the best ORBW is 0.5 nm. The influence of the change of ORBW from 1 nm to

7.5

5.5

60£r

± ± ±

-40 -30 -20 -10

Pin [dBm]

Figure 1: Noise figure as function of input power for 10, 25 and 60 degr. C

0.5 nm on the noise figure can be seen in trace B

in figure 2. This trace is positioned around 0.4 dB

higher than trace A as expected from the decrease in ORBW. Trace A and B were measured both using

input power measured after the source connector and

the EDFA output connector in open air. The power

calibration was done by a power meter using a lensed

photodetector. Changing to the definition of input

power to fibre guided power and using a large area

detector without any lenses led to trace C in figure 2. Trace C in figure 2 will be used as a reference line in the discussion that follows. Five different labs up to now have mesisured the

RREDFAs noise figure at a wavelength closely around

1555 nm. One lab used both the OSA (I) and ESA

method (lA), three only the OSA method (II, III and

V) and one merely the ESA method (IV). We have

plotted the results in figure 3. Two following observations can be made from figure 3:

• Labs I and III use identical procedures in cali- bration and definitions. The noise figures they

get are within a range of 0.3 dB, which is quite

good.

• Lab II, the manufacturer, obviously is rather

optimistic on its noise figure, this is most probably caused by a bad ORBW and absolute optical power calibration

• Lab V makes a large error in its optical bandwidth calibration which results in a large deviation from the rest of the data.

12
-50 -40 -30 -20 -10 0 Pin [dBm]

Figure 2: Noise figure measured by OSA method as a function of input power. Trace A first measurement;

Trace B: second measurement after calibration of ORBW;

trace C: third measurement after changing definitions of input and output power and using a large area detector

for calibration of absolute optical power.

-40 -30 -20 -10 Pin [dBm]

Figure 3: Noise figure as a function of input power as measured in five different labs. See text for explanation

of legend.

• Lab IV uses the measurement of the carrier to noise ratio with an analog transmission system.

The RIN of the laser is not low enough to be

discarded, and input and output shot noise levels are different. This means that complicated

corrections must be made resulting in data that

are near the traces from labs I and II for the

lower input powers, but deviate significantly for higher input powers, showing that conversion

of the electrical data to the "right definition" is intricate.

• The measurement at lab I with the electrical method (lA), shows good agreement with the

values obtained by its OSA measurement (trace

I) for an input power of -5 dB. However for higher input powers the noise figure increases

steeper. Whether this is caused by the fact that

"inversion noise" start to play a role, as claimed

in section 3 or simply by measurement error, remains to be investigated.

5 Items to be specified in the

measurement protocol

The accuracy of the measurement depends of course on the accuracy of the equipment that is used. This

accuracy is in turn is determined by proper calibra- tion by the user itself, and the "intrinsic" accuracy

of the system. The latter is related e.g. to scale fi- delity, dynamic range of the equipment and cannot

be controlled by the user in most cases. Next we will discuss the items that can and should be influenced

by the operator of the measurement setup.

• Connector: The best way to avoid problems

with connector losses is to avoid connectors by

fusion splicing. However this requires a fusion

splicer at every lab. Therefore the second best

option is the use of a low-reflection connector. These can not be handled by every lab

either. The approach we choose is that the co- ordinating lab receives the RREDFA back af- ter each measurement, and supplies it by fusion

splicing with the connector appropriate for the next round-robin candidate. The coordinating

lab can characterize the connectors and thus

track possible measurement differences caused

by connector losses.

• Input and output power location: The power

measured by a power meter is not the fiber guided power, but the fraction that leaves the

fiber after the glass- air surface. The power entering the EDFA however is the fibre guided

13
power, minus connector losses of course. The

differences between those two power levels are around 0.16 dB and influence directly the out- come of a noise figure measurement.

• For ESA method: minimum input power: To

ensure that the signal spontaneous beatnoise

dominates. If the input power in the EDFA

drops below a certain level, sp-sp beatnoise

starts to influence the total noise level. Al- though it is in principle possible to include this

effect in the conversion equations, it adds an er- ror source and should therefore be omitted for a round-robin.

• For ESA method: RIN of the transmitter laser:

If the noise of the transmitter laser is not shot

noise limited, which is usually the case when

the laser's intrinsic RIN exceeds -155 dB/Hz, we need to know this RIN to correct for it in the calculations.

• Optical Bandwidth Calibration: The optical

bandwidth directly enters the conversion equation 2. The bandwidth readout at the optical

spectrum analyzer can not be trusted and is usually not the same over the whole wavelength

range. Often there is one optimum bandwidth,

where the calibration is optimal. It is prefered to do a careful calibration using a very narrow

linewidth tunable laser.

• Absolute power calibration: The measured ASE

level enters directly the conversion equation 2 and depends on the absolute power level cali- bration, which is usually done by a calibrated

optical power meter. Differences can be caused

by using a large area detector i.s.o. a lens cou- pled small area device.

• Subtraction of source noise: The amplified source noise can not be distinguished from the

ASE noise power and should be subtracted from

the measured ASE, either by calculation or by

making use of the polarization properties of the source light.

6 Conclusions

Besides the fundamental diff"erences between the ESA

and OSA noise figure measurement method caused

by the fact that photon noise statistics at high input

powers differs from the assumed statistics and caused

by the inability of the OSA method to measure inter- ferometric noise, the more "trivial" problems of measuring an EDFA's NF properly have been discussed.

This discussion was based on observations made during a "mini" round-robin along five European laboratories up to now. Most of the diff'erences between

the results can be attributed to using calibration of

the optical resolution bandwidth and absolute optical

power as far as OSA is concerned. Major problems

with the ESA method are in the proper subtraction

of source noise and handling diff'erent shot noise levels. Both the ESA and OSA method require clear definition of the positions where the optical power is measured.

Acknowledgements

The following individuals provided measurements:

Chris van der Plaats, AT&T Network Systems

Netherlands; Martin Kowatsch, Alcatel-SEL, Andreas Klar, Universitat Kaiserslautern; Edgar Leckel,

Hewlett-Packard Boeblingen. The discussions with

Christian Hentschel and Edgar Leckel from HP were

highly appreciated. This work was partly supported

by the Commission of the European Union under the

RACE II program, project R2062, "Components for Fibre to the Home (COMFORT)".

References

[l] E. Desurvire, "Analysis of noise figure spectral distri- bution in Erbium doped fiber amplifiers near 980 and

1480 nm.," Appl. Opt., vol. 29, pp. 3118-3125, July

1990.

[2] F. W. Willems, J. C. van der Plaats, M. A.

Jutte, and M. Babeliowsky, "Comparison of satu- rated EDFA noise figure determination with ESA and

OSA method," in Proc. of. Optical Fibre Measurement

Conference, OFMC'93, Turin, Sept. 1993.

[3] F. W. Willems, J. C. van der Plaats, and D. J. DiGiovanni, "EDFA noise figure degradation caused

by amplified signal double Rayleigh scattering in Er- bium doped fibres," lEE Electronics Letters, vol. 30, pp. 645-646, Apr. 1994.

[4] E. Jaunart, P. Crahay, and M. Blondel, "EDFA

noise figure compuatation for small and large signals

regime.," in Digest of COST 94 Workshop Optical

Telecommunicationsm, Nice, April 1994-, pp- 264-

267, Apr. 1994.

[5] E. Desurvire, Erbium- doped fiber amplifiers. John Wiley & Sons, 1994.

[6] J. AspeU, J. F. Federici, B. M. Nyman, D. L. Wilson,

and D. S. Shenk, "Accurate noise figure measurements

of erbium doped fiber amplifiers in saturation conditions," in Digest of Conference on Optical Fiber Communications '92, San Jose, pp. 189-190, Feb. 1992. paper ThA4.

14
Crosstalk in 1.3pm. Praseodymium Fluoride Fibre Amplifiers

S.C.Fleming

Optical Fibre Technology Centre, University of Sydney, Australia^

Abstract

Praseodymium doped fluoride fibre amplifiers have been demonstrated to be practical components for 1.3|im. telecommunications systems. This paper presents the first experimental measurement of the

crosstalk performance of a Pr^:fluoride fibre amplifier. An experimental setup is described in detail. Measurement results are presented and compared with the results of a theoretical analysis. Good

agreement between measurement and theory is demonstrated. Conclusions are presented concerning the

implications of the measured crosstalk performance of these amplifiers for real systems. Furthermore, the

first experimental evaluation of the lower state lifetime of the Pr^'':fluoride system is derived from these

measurements.

Introduction

Recently there has been much interest in praseodymium doped fluoride fibre amplifiers for use in the

second telecommunications window. Gains of over 30dB [1] and saturated output powers over 200mW [2] have been demonstrated. More recently Nd:YLF lasers have been demonstrated as practical pump sources

for efficient Pr^:fluoride amplifiers [3]. With excellent levels of gain and saturated output power

demonstrated using practical pump sources there is now a need to examine other characteristics of these

amplifiers, of importance to their application in real telecommunications systems.

Crosstalk is the transfer of modulation or information

between two signals. This paper addresses crosstalk

between two such signals, both within the gain band of the

amplifier. This is an important parameter for amplifiers

where there will be more than one signal within the gain

band. Given the broad and relatively flat gain band of Pr^:fluoride fibre amplifiers they will be especially attractive

for use in wavelength division multiplex systems where

potentially many signals at slightly different wavelengths

around 1 .3|im. will be used simultaneously.

Figure 1 shows the energy level diagram for Pr'*. Pumping

at around 1|im. creates an inversion between levels 'G4

and ^Hg producing gain for signals around 1 .3|im. Crosstalk

involves the modulation of the inversion and hence the

gain. A large signal around 1 .3|im. in the amplifier will tend

to reduce the inversion and hence the gain available to any

other signals. If the large signal is modulated, the inversion

and gain will pick up this modulation, approximately in anti-phase. Any other signals around 1.3|j.m. will thus

experience a time varying gain and be modulated correspondingly.

Pump Signal

rNonRadiative

r

1G4

3F4

3F3

3F2

-3H6

3H5

3H4

Figure 1 - Pr^* Energy Levels

^ This work was carried whilst the author was with the Networks Research Division, BT Labs,

Ipswich, UK.

15
Experiment

Figure 2 shov;s a schematic diagram of the experimental setup used in this worl<. An amplifier fibre consisting of 7m of -0.2 NA, SOOppm Pr^*:fluoride fibre was pumped with around 1W at lOIOnm from a

Ti:Sapphire laser through a wavelength division multiplexer (WDM). 7m is not the optimum length for such an amplifier and hence the pumped/unpumped change in transmission was around 22dB (giving an

input/output gain of around 18dB). Unabsorbed pump light was blocked by a silicon filter in a beam

expander. A pump block is essential since this high level (compared with the signals) of remnant pump has

also picked up modulation since the pump absorption increases when the large signal pulls down the

inversion. This pump block was later replaced with a high concentration ytterbium doped silica fibre which was found to be a better absorber for the remnant pump with lower losses at the signal wavelengths.

Praseodymium fibre

TiiSapphire Laser

Tunable

Optical Filter

Isolator

Figure 2 - Experiment for Measuring Crosstalk

A ~100|iW, 1 .31|im probe signal, from a tunable long external cavity laser, was launched, co-propagating

with the pump, through the WDM. A 2mW, 1.29|im distributed feedback laser was modulated, from 5Hz

to 200MHz, by a network analyser. The signal from this laser diode was launched via a coupler into the

amplifier, counter-propagating with respect to the other two signals. Having the probe and perturbing

signals counter-propagating significantly complicates the analysis of the transfer of modulation. However

it also significantly simplifies the experiment in that it is much easier to separate these two signals (which

are at very similar wavelengths and carrying the same frequency modulation) if they are not propagating

in the same direction. It was necessary to use a very unbalanced coupler (-95/5%) in order to maintain a

sufficiently high level of modulated signal within the amplifier to perturb the gain.

This large modulated signal reduced the gain, by around 1dB at the peak, resulting in a modulation transfer

to the amplified probe signal. After passing through the pump-block filter, -5% of the amplified probe signal

is coupled off and passes through a tunable optical filter and onto a detector. Any back-reflected perturbing

signal, amplifier ASE and remnant pump is blocked by the tunable optical filter. By tuning the optical filter and the probe laser, the spectral separation of the probe and perturbing signal could be varied and, in principle, any variation in crosstalk with spectral separation, due to inhomogenous broadening, could be

measured.

16
A network analyser was used to process the amplified detector signal to give the amplitude and phase of the transfer of modulation from output to input. To calibrate the measurement system it was necessary also

to measure the transfer functions of the electrical and electro-optical components and the finite length of

fibre between the modulated laser diode, the detector and the amplifier input. This was performed by

breaking the connection between the pump-block output and the amplifier input and tuning the optical filter to the perturbing signal wavelength. The -4% reflection from the bare end of the pump block pigtail fibre could then be used to measure the response of these other components and hence to calibrate a crosstalk

measurement.

Results and Conclusions

To analyse the measurement results it was necessary to develop a model for the crosstalk performance

of a Pr^:fluoride fibre amplifier. Development of an exact analytic description of this is almost certainly an

intractable problem. However, a perturbation analysis [4] of the rate equations yielded a set of approximate

analytic equations describing the transfer of modulation to the inversion from a small modulated signal. A

range of parameters were used in the evaluation of these equations and the results compared with those

of an exact dynamic rate equation model. It was found that the perturbation analysis gives simple

approximate analytic equations which agree well with the exact solution.

A full propagation model was written incorporating these perturbation equations. Using this model the

steady state solutions for a pump and the mean levels of a perturbing signal and a non-perturbing probe

signal throughout a fibre amplifier were calculated. Typical cross-section and lifetime values were used

(-C2i=100nS, T32=1 10m,S, ap=4.00x10'^m^, as=3.48x10'^^m^). Exact field shapes were taken into account but

not amplified spontaneous emission. Steady state values of the pump and signal rates and the inversion

at each point in the fibre were calculated and used to evaluate the analytic equations derived from the

perturbation analysis. From this evaluation of the amplitude and phase of the crosstalk at each point in the

fibre the overall crosstalk of the amplifier was calculated, taking into account whether the probe and

perturbing signal were co- or counter-propagating with respect to each other.

Figure 3 shows a comparison of the measured and calculated results. Parameters matching those of the

experiment were used to run the model. As can be seen there is excellent agreement, in both amplitude

and phase, between the two data sets from low frequency to around 4MHz. This upper frequency limit was

set by the measurement noise floor; at this frequency the crosstalk had decreased to a level where it could no longer be extracted from the noise. Extension of the measurement range to higher frequencies could

be simply achieved by either increasing the perturbing signal level or the amplifier gain.

The general trends are as expected with the crosstalk amplitude essentially constant at low frequencies,

rolling off at a frequency in the region of the inverse of the upper state lifetime (the exact position of this breakpoint is a strong function of the pump rate) and decreasing at lOdB/decade at higher frequencies.

The crosstalk phase response starts in exact anti-phase at low frequencies moving to a Td2 phase lead

at higher frequencies.

Some specific features of this response are of interest. At high frequencies on the phase plot the roll-off

is due to the probe and perturbing signal counter-propagating. In contrast, it is found that in the co- propagating case the high frequency phase response tends to a 7i/2 phase lead. Despite this significant

difference in the phase response, there is negligible difference in the amplitude response between the co- and counter-propagating cases.

At around 1 MHz there is a distinct feature in the theoretical responses which is due to the lower state

lifetime but occurs a decade lower in frequency than the inverse lifetime. Inspection of the measured

response shows that this feature is not present for frequencies below 1MHz and is not apparent, above

the noise, at higher frequencies. Hence an upper limit of around 50nS can be placed on this, previously

unmeasured, lifetime.

17
Figure 3 - Calculated and Measured Crosstalk

Conclusion

The general form of the crosstalk is very similar to that of an erbium doped fibre amplifier [5]. Hence,

probably the most important conclusion of this work is that there are unlikely to be any new problems

encountered with crosstalk in Pr'^ifluoride amplifiers. Specifically, for digital systems operating at hundreds

of MHz or more the crosstalk will be negligible. However as for erbium amplifiers problems may be

encountered with low frequency supervisory signals or with analog systems.

It is also of significance that the lower state lifetime has been found to be at least four orders of magnitude

smaller than the upper state lifetime. This indicates that there will be no significant problems with

Pr^*:fluoride fibre amplifiers due to "bottlenecking" [6] and that the effect that this level has on the crosstalk

performance is negligible. This lifetime has not been previously measured since it is a non-radiative

transition and not measurable by fluorescence decay.

Acknowledgments

The author gratefully acknowledges Richard Wyatt and Tim Whitley for many useful discussions. This work was funded in part by the Commission of the European Community, Project R2018.

References

1 - Y.Ohishi et al, IEEE Photonics Tech. Lett., 3, 8, pp.71 5-71 7 (Aug. '91)

2 - T.J.Whitley, OSA OAA'92, p8 (1992)

3 - T.J.Whitley et al, IEEE Photonics Tech. Lett., 4, 4, pp. 399-401 (Apr. '93)

4 - T.J.Whitley, private communication (Oct. '91) . 5 - E.Desurvire, IEEE Photonics Tech. Lett., 1, 8, pp.196-199 (Aug. '89) .^ v 6 - W.P.Urquhart, IEEE J. Quantum Elect., 28, 10, pp.1 962-1 965 (Oct. '92)

18
MEASUREMENT AND COMPENSATION OF GAIN-PEAKING IN MULTI-AMPLIFIER LINKS

Evan L. Goldstein

Bellcore, Red Bank, NJ

ABSTRACT

The emergence of practical fiber- amplifier chains has raised

the prospect of transparent lightwave networks, in which signals travel from source to destination through a sequence of

intermediate nodes without optoelectronic conversion. When

such networks employ multiple wavelengths, however, some of

the most substantial new research challenges are those posed

by the amplifier chains themselves. Such networks suffer from

accumulating inter- channel power-spread, from sensitivity to inter-amplifier loss variations, and from transient cross- saturation, as the network undergoes reconfiguration. We here describe two different approaches to address these problems.

I. Introduction

Due to the superior noise performance, erbium-doped fiber- amplifiers (EDFAs) have demonstrated in single-wavelength

point-to-point communications links, there is growing interest in employing EDFA cascades in reconfigurable multiwavelength networks that can be scaled to large size [1]. The

chief obstacle facing such cascades arises Irom amplifier gain

nonuniformities. Each amplifier provides gain, in dB, given

by GdB(A) ~ /Q[cre(A)n2(z) - cJa{X)nx{z)\dz, referred to below as the spectral gain equation. Here <7e(A) and cra(A) are the emission and absorption cross-sections of the erbium ions,

ni(z) and 712(2) are their lower- and upper-state linear population densities at position z along the amplifying fiber, and

/ is the ampHfier's length [2]. The integrand of the spectral

gain equation exhibits a wavelength-dependence imposed by

o-e(A) and (Ta(A), which are in turn determined by the spectroscopic properties of erbium atoms in silicate glass. While this wavelength-dependence can be modified, via ni(z) and ^2(2;), by adjusting pump power, input signal power, erbium-ion

density, and fiber length, it cannot be eliminated. This implies that wavelength-multiplexed channels traversing a chain

of nominally identical amplifiers will develop an inter-channel

power spread that grows exponentially along the chain. This

effect is illustrated in Fig. 1, showing the output spectrum

after the first and last amplifier in a six-amplifier cascade.

The five channels are equalized at the input of the cascade.

After one amplifier a power spread of ~ 3 dB is seen. The

spread grows exponentially along the cascade and has reached

1520 1530 1540 1550 1560 1570 1580

Wavelength [nm]

Figure I: Inter-channel power spread accumulates through

amplifier cascades.

~ 16 dB after the sixth amplifier. Thus, as a multiwavelength

network is scaled up in size, channels residing off the gain peak

will fall toward power levels that are undetectable in the pres- ence of receiver noise or amplifier-induced beat noise or both.

Furthermore, there are significant additional challenges. In order to build amplifier chains that will suffice for large

wavelength-division-multiplexed (WDM) communications net- works, two additional problems must be surmounted. Like the

problem of accumulating inter-channel power-spread articu- lated above, both of these are inherent in the spectral gain

equation, and are thus more or less fundamental. The first problem is that variations in inter-amplifier loss result, via

saturation, in variations of ni{z) and n2{z) in the system's

amplifiers and consequently in the gain spectra GdB(A) of the

individual amplifiers. These in turn accumulate to produce

variations in the accumulated gain spectrum of the chain.

Thus, WDM amplifier chains are not in general particularly

robust in the presence of inter- amplifier loss-variations [3]. In single-wavelength cascades, this problem is elegantly addressed by operating the amplifiers at moderate-to-high levels 19
10

ffi -10

k.. »

I

-20

-30

-40

AMPLIFIER 6

O 1 CHANNEL

3 CHANNELS

5 CHANNELS IP

Uu

Combiner -10 dB

1520 1530 1540 1550 1560 1570 1580 Wavelength [nm]

Figure 2: Dynamic cross-saturation arises when WDM

channels are turned on or rerouted in optical networks with

amplifiers.

of gain-compression, which provides automatic compensation

for inter- amplifier loss-variations [4] [5]. However, in multiwavelength cascades, this tactic loses all of its effectiveness,

since it serves to regulate only total optical power, not per- channel power.

The second problem is that when a transparent

WDM communications system undergoes network reconfiguration, e.g. by adding or rerouting just one channel, the added

power in this channel will stimulate emission and absorption

in the first amplifier it encounters, altering ni(z) and n2{z)

and thus, through (1), the amplifier's gain spectrum. Once

again, these perturbations will accumulate along the chain

of amplifiers. The result is that by adding or rerouting one

channel in a reconfiguring WDM network, one perturbs other

channels sharing all or part of the route. This is illustrated in

Fig. 2 where the peak power levels of the individual channels

are shown at the output of the six-amplifier cascade with one, three or five channels present at the input. The output power

variation of the 1546 nm channel is 16 dB when the loading

changes from just one channel to all five channels.

In the following sections, we discuss two different approaches to address these problems: cooled inhomogeneously broadened amplifiers and pump-shared amplifiers in parallel.

II. Accumulating inter-channel power-variations

When wavelength-multiplexed signals traverse a single fiber amplifier, the channels experience spectral gain variations

that are generally modest. The serious problem, however, is that these variations accumulate from stage to stage, rapidly

growing far too large to allow adequate noise performance at the less-favored signal wavelengths. Efforts to solve this prob10 %Rbef Coupler

Dichroic Coupler

6-cm

Fiber Coll InUquklNi

Figure 3: Experimental setup for investigating inhomogeneously broadened amplifiers.

lem have thus far centered on two approaches. In moderatesized point-to-point systems, boosting the transmitter power

for channels with low gain [6] is quite effective. However, this

will not generally suffice in multipoint-to-multipoint light- wave networks (e.g. [7]) with dynamically varying signal paths, or in ultra-long-haul point-to-point links. One can also use

various optical-filtering schemes to selectively attenuate highgain channels [8] [9]. However, in such schemes, inter-channel

power variations still accumulate from stage to stage unless a

separate active servo-loop is used for each wavelength.

We here show that when cooled, gain-saturated fiber amplifiers are placed in cascade, their gain is effectively clamped on a channel-by-channel basis by the propagating signals themselves, without the use of feedback arrangements or servo- loops [10]. The system thus behaves as a parallel overlay

of independent amplifiers for the various channels. Further- more, we show the effectiveness of using physically separated

pump-shared amplifiers in parallel as a room-temperature implementation of per-channel amplification.

II. A Cooled inhomogeneously broadened amplifiers

Experiment. The experimental arrangement, shown in Fig. 3, consists of a cascade of ten erbium-doped aluminosilicate fiber amphfiers. Each is codirectionally pumped with ~11 dBm of

light at 1480 nm giving a small-signal gain of ~ 14 dB. The

20-m-long doped fibers were wound into coils of 6-cm diameter, so they could be immersed in liquid nitrogen at ~77 K.

Each fiber amplifier is separated from its successor by a trans- mission fiber with about 10 dB of loss. The transmission

path contains neither optical filters nor, apart from the inhomogeneous amplifiers themselves, any special channel-powerequalizing components.

Results and Discussion. Five wavelength-multiplexed channels ranging from 1546 nm to 1560 nm, at —15 dBm per

20
channel, were coupled into the first amplifier. Inter-channel

wavelength separations were about 4 nm, which is considerably larger than the measured ~l-nm homogeneous linewidth

of the amplifiers at 77 K [11]. We measured the optical power

spectra emerging from each amplifier stage under two conditions: with aU amplifiers at room temperature (293 K), and

with all at 77 K.

The plots in Fig. 1 show the output power of amplifiers

1 and 6 operating at room temperature. After amplifier 1, the signals vary in power by about 3 dB. However, after amplifier 6 the variation has accumulated to about 16 dB, with

the weakest channel, at 1546 nm, transporting only about

-18 dBm. Most importantly, the inter-channel power varia- tions are accumulating rapidly as the signals pass amplifiers

1 through 6. The trends are more clearly seen in Fig. 4, which plots

the output power of the five wavelength-multiplexed signals versus amplifier number for the full 10-amplifier cascade. At room temperature (Fig. 4a), the inter-channel power spread

accumulates at the rate of about 3 dB per stage, and totals 27 dB after amplifier 10. At 77 K (Fig. 4b), the signal powers spread out slightly in the first few amplifiers to about 5 dB by ampUfier 4, after which the accumulation

slows greatly. The accumulation is completely arrested, up

to uncertainties imposed by experimental error, in amplifiers

7, 8, 9, and 10. It is worth noting that the cooled, inhomogeneous amplifier's saturation-output power is remarkably

high, in view of the fact that each channel is coupled to only

a subset of the fiber's erbium ions. Indeed, after the sixth

amplifier, the weakest channel (1560 nm) still transports -5 dBm of power. Thus, in the inhomogeneovis chain the sig- nal powers not only converge toward a fixed or nearly fixed

interchannel power spread; they also propagate at readily detectable power levels. In conventional, homogeneously broadened amplifier chains, whether saturated or not, the inter- channel power-spread monotonically accumulates from stage

to stage, thus hmiting system scalability. In a saturated, gaindecoupled chain, however, e.g. the above inhomogeneously

broadened system, the per-channel powers asymptotically ap- proach spatially fixed values. This provides automatic per- channel power-regulation, and prevents the spatially accumulating inter-channel power spread that prevents transparent

WDM lightwave-communications networks from being scaled

up in size.

II.B Pump-shared parallel amplifiers

We here show that the uncomplicated approach of placing

pumi>shared parallel amplifiers between a multiplexer pair

2 4 6 8 10

Amplifier Number

12

Figure 4: Per-channel output power vs. amplifier number

for: a) Room temperature (293K) homogeneously broadened and b) cooled (77K) inhomogeneously broadened amplifiers.

stops the accumulating power-spread, so long as the amplifiers are properly configured. Moreover, if properly configured, no

per-channel attenuators or servo-loops are required. Systems

constructed from such modules self-adjust for transparency at each signal wavelength without servo-loops since the channels

effectively have been gain-decoupled in the amplifier chain.

Given a large inter-channel power spread, they will compress

the spread in each amplifier module, rather than accentuating it. Although it is not widely appreciated that properly

configured systems have these properties, the approach itself

is based on well-known saturation phenomena. What is novel

is the relative simplicity with which a fundamental problem,

inherent in the gain dynamics of fiber amplifiers, can be addressed. Experiment. To demonstrate the properties summarized

above, we have performed measurements on the amplifier

module shown in Fig. 5. Four twenty-meter lengths of erbiumdoped fiber share a 70-mW 1480-nm pump laser, with each

fiber receiving ~ 13 mW of pump light, and providing ~ 14 dB

of small-signal gain at 1550 nm. The fibers were placed between interference-filter-based wavelength multiplexers having nearly fiat, 1.5 nm passbands at the signal locations of

1546, 1550, 1554, and 1558 nm. Each multiplexer imposed 1.5 dB of insertion loss on each channel, and 27 dB of isolation between any pair of channels. Each dichroic coupler added

~1 dB of insertion loss at 1550 nm. Thus, when the assembled amplifier module was operated single-stage (i.e. without

21
Er-doped

Fiber X-Mux

1

Signal

in First Gain Stage

(Optional)

DC

-n- -a D.

D.

Splitter

A-Mux

Signal Out

I Laser

Figure 5: Elxperimental setup. DC: dichroic coupler. A- Mux: interference-filter-based wavelength-multiplexer.

E

CD

•a«

o

a.s

a. o

-40

- 30 - 20 - 1 0 Input Power [dBm]

1 0

Figure 6: Per-channel input power versus per-channel out- put power for the single-stage and dual-stage amplifier

modules, measured with all per-channel powers varied in unison.

the first gain stage), it provided each channel with a smallsignal gain of 10.0 ± 0.2 dB. The optional first gain stage

in Fig. 5 consisted of a conventional, bidirectionally pumped,

single-stage fiber amplifier employing 40-mW, 1480-nm pump

lasers and operated at a relatively strong inversion level of

n2/ni ~ 3 [12]. Its small-signal gain at 1550 nm was 25 dB.

Results and Discussion. Figure 6 shows the measured

output power versus input power at each of the four channel

locations for the single-stage amplifier module, and for the

dual-stage module employing the optional first gain stage.

Power levels were measured at the positions marked 'signal

in' and 'signal out' in Fig. 5. At low input-power levels, little inter-channel power-spread is seen in Fig. 6. At high in- put power, the output-power spread approaches 0.8 dB in the single-stage module, and 1.5 dB in the dual-stage module. Power spreads of this magnitude are readily tolerated in communications systems. The question we wish to focus on, however, is that of 'scalability': does the inter-channel power

spread accumulate from stage to stage?

The chief virtue of this system is that inter-channel powerspread does not accumulate from stage to stage, so long as the system is operated in gain-compression. On the contrary,

when illuminated with signals at widely divergent per-channel

powers, the system will tend to compress the inter-channel

power-spread to a small, fixed value. We show this through

measurements on a single amplifier module, since we were

unable to perform measurements on cascades of them.

The performance of a cascade of single-stage modules is easily predicted to reasonable accuracy by using the fact that

the per-channel gains are decoupled. One need only choose a

value of the per-channel power P\n incident on each amplifier; this design choice will in turn be constrained by the accu- mulation of degradations due to spontaneous-emission noise

and fiber nonlinearity. The target loss L between amplifier

modules is then given implicitly by the expression

LGo exp [(1 - L-i)(Pi„/Psat)] = 1, (1)

obtained from Ref. [4] by setting the saturated gains of the

modules equal to the losses between them. Here Go is the

amplifier module's small-signal gain and Pgat is its saturation

power, with all quantities in linear units and L < 1. Eq. (1) has a nonzero solution for Pin, provided that Go > L~^; thus,

the amplifiers must be operated in gain-compression. If L

varies along the chain, Pjn will also show variations, though

these will be partly compensated by the compressed amplifiers' self-adjusting gains [4]. The single-stage curve of Fig. 6 in fact represents a measurement of the relations implicit in (2) if one sets L =

Pin /Pout- Thus, choosing a reasonable input-power level of

—10 dBm, Fig. 6 shows that an inter-amplifier loss of 8 dB

(= —2dBm + lOdBm) should be selected. Given a set of per- channel powers incident on a cascade of such modules, their evolution along it can then be computed through iterative use

of Fig. 6. An example is plotted in Fig. 7, where we have taken

[^1546, ^l550> ^{554,^1558] = [-20,-15,-5,0] dBm, with 8 dB of loss between amplifiers. Although the amplifiers are gain-compressed by only ~ 2 dB, it is seen from Fig. 7 that

the system rapidly power-equalizes. The inter-channel power

spread, initially 20-dB, drops to ~ 10 dB at the input to the

third module, thereafter approaching an equilibrium value of ~ 1 dB. The spread then remains fixed from stage to stage.

III. Transient cross-saturation

The last problem we address, transient cross-saturation, is unique to transparent amplified WDM networks that undergo

physical network reconfiguration during normal system operation. It does not arise in point-to-point links. Although

22
- 5

1 0

1 5

20

25

.....

J

-—d^—

1 1r "- T 1 I III.

# 1546 nm ; -O-1550 nm j --1554nm ; -Q-1558nm 1

- 1

i i i i i i i i i 3 4 5 6 7 8 9 Amplifier Module Number

10 11

E

CDo

10

0 -

-10

o$ -20 o

Q.

-30

-40

77K

O 1 CHANNEL

3 CHANNELS

5 CHANNELS

AMPLIFIERS

1520 1530 1540 1550 1560 1570 1580 Wavelength [nm]

Figure 7: Calculated evolution of the per-channel powerlevels along a cascade of single-stage modules.

Figure 8: Transient cross-saturation in a cascade of six in- homogeneously broadened amplifiers.

transient cross-saturation in fiber amplifiers is a well-known

phenomenon [13], it is widely believed that WDM systems

employing fiber amplifiers are, in practice, immune from it due to the device's slow saturation dynamics. As has long

been known [13], the saturation process in fiber amplifiers is characterized by a time constant of ~ 3(X)/2s; consequently,

there is effectively no transient cross-saturation at frequency

components higher than ~ 5 kHz. This does indeed imply

that ampUfied point-to-point high-speed WDM transmission

links suffer no ill-effects from transient cross-saturation. However, amplified WDM lightwave networks are not immune if they undergo reconfiguration with Fourier components at kHz

frequencies and below. This is in practice precisely where

those Fourier components are likely to be, due to the slow,

computation-intensive process of determining and broadcasting the network switch settings required for a given network

reconfiguration. Thus, one must build into amplified multiwavelength networks some mechanism which confers immunity to transient cross-saturation.

Clearly, inhomogeneous broadening, or more generally the

decoupling of per-channel gain dynamics, provides such a mechanism. It is because of the homogeneous broadening of

conventional fiber amplifiers that cross-saturation will arise

in networks. If the amphfiers are instead inhomogeneously

broadened, they will saturate on a channel-by-channel basis,

leaving their signals unperturbed by turn-ons and reroutings

at other wavelengths. Thus, inhomogeneous amplifiers not

only help power-equalize a wavelength-multiplexed network;

they also suppress the dynamic inter-channel cross-saturation

that occurs during network reconfigurations [14].

This beneficial crosstalk suppression was again experimentally measured using the set-up of Fig. 3. Signal light from

five wavelength-multiplexed lasers, transporting -15 dBm per

channel, was sent into the first amplifier. Various signal channels were then switched on and off, simulating multi-access network reconfigurations. The signal spectrum was monitored at the output of each amplifier in the chain using the

10% probe couplers of Fig. 3. Figure 2 plots data taken for three system configurations: single-channel (1546 nm only),

three-channel (1546, 1554, 1560 nm), and five-channel (1546,

1550, 1554, 1557, 1560 nm).

Figure 2 shows that when the homogeneous system is op- erated single-channel, reconfiguring to three channels results in a 12-dB drop in signal power at 1546 nm. Reconfiguring to five channels produces a 15-dB drop. This is due to homogeneous inter-channel cross-saturation in the cascaded

amplifiers. The perturbations are enhanced by the fact that

the 1546-nm channel lies well off the gain peak of the roomtemperature amplifiers.

In the cooled, inhomogeneous system, by contrast, a power

drop of only 2 dB occurs when reconfiguring to three channels, as shown in Fig. 8. A drop of 4 dB occurs when re- configuring to five channels. Thus, when moving from singlechannel to five-channel operation, the inhomogeneous amplifiers suppress inter-channel cross-saturation by about 11 dB.

The improvement results from the decoupling of wavelengths

in saturated, inhomogeneous, cascades. However, some homogeneous broadening is evident in the cooled system too; as a result, the single-channel and three-channel points in Fig. 8 do not lie precisely at the peaks of the five-channel spectrum.

As for the pump-shared parallel amplifiers, the singlestage amplifier showed complete decoupling of saturation dynamics for the various channels, as one would expect given

3 J
the multiplexer parameters stated above; its per-channel gain

and output-power levels thus showed no measurable sensitivity to the number of channels incident on the module. The

dual-stage module showed slight coupling, as one would also expect, due to gain-competition in the first stage.

rV. Summary

To summarize, we have seen that fiber-amplifier chains for re- configuring WDM communications networks pose not simply

the problem of the non-flat gain-spectrum, difficult though

that problem in fact is. They instead pose a set of related

challenges, all of which result from the non-flat gain-spectrum

of erbium atoms in silicate glass, together with the homogeneous saturation properties inherent in the spectral gain

equation of section 1. Transparent amplified WDM networks

using conventional fiber amplifiers exhibit spatially accumulating inter-channel power-spread and sensitivity to changes

in inter- amplifier loss. Moreover, when such systems un- dergo dynamic reconfiguration, they are vulnerable to transient inter-channel cross-saturation. These properties are in- herent in the spectroscopy of the gain medium, and are thus

reasonably thought of as fundamental.

Nonetheless, we have also seen that all three challenges can be addressed in one stroke if the per-channel gain dynamics of the amplifier modules can be decoupled, e.g. by

employing an inhomogeneously broadened gain medium or by amplifying each channel in a physically separate amplifier. Saturation can then be employed to yield a system that

power-regulates on a per-channel basis, rather than on a total- power basis. The result is an automatically equalizing system

that not only arrests the inter-channel power-spread but also provides tolerance to inter-amplifier loss variations as well as reduces the transient cross-saturation present in transparent

reconfigurable optical networks.

Acknowledgement. This work was partially funded by the

Advanced Research Projects Agency (ARPA) under the Optical Networks Technology Consortium.

[4] C. R. Giles and E. Desurvire, "Propagation of signal and

noise in concatenated erbium-doped fiber optical amplifiers," IEEE J. Lightwave Technoi, vol. 9, pp. 147-154,

1991.

[5] N. S. Bergano, J. Aspell, C. R. Davidson, P. R. Triscitta, B. M. Nyman, and F. W. Kerfoot, "A 9,000 km 5 Gb/s

and 21,000 km 2.4 Gb/s feasibility demonstration of

transoceanic EDFA systems using a circulating loop," Proc. Conf. Optical Fiber Commun., 1991, paper PDP13.

[6] A. R. Chraplyvy, J. A. Nagel, and R. W. Tkach, "Equalization in amplified WDM lightwave transmission sys- tems," IEEE Photon. Technoi. LeU., pp.920-922, 1992.

[7] C. A. Brackett, "The principles of scalability and modularity in multiwavelength optical networks," Proc. Conf.

Optical Fiber Commun., paper TuJ2, 1993.

[8] K. Inoue, T. Kominato, and H. Toba, "Tunable gainequalization using a Mach-Zehnder optical filter in multistage amplifiers," IEEE Photon. Technoi. Lett, pp. 718-

720, 1991.

[9] S. F. Su, R. Olshansky, G. Joyce, D. A. Smith, and J. E. Baran, "Gain-equalization in multiwavelength light- wave systems using acoutooptical tunable filters," IEEE

Photon. Technoi. Lett., pp.269-271, 1992.

[10] E. L. Goldstein, V. da Silva, L. Eskildsen, M. Andrejco,

and Y. Silberberg, "Inhomogeneously broadened fiber- amplifier cascade for wavelength-multiplexed systems,"

IEEE Photon. Technoi. Letts., vol. 5, pp. 543-545, 1993.

[11] E. Desurvire, J. L. Zyskind, and J. R. Simpson,, IEEE

Photon. Technoi. Lett., pp. 246-248, 1990.

[12] E. L. Goldstein, L. Eskildsen, C. Lin, and R. E. Tench,

"Multiwavelength propagation in lightwave systems with

strongly inverted fiber amplifiers," IEEE Photon. Technoi. Lett, vol. 6, 1994.

[13] C. R. Giles, E. Desurvire, and J. R. Simpson,, "Transient

gain and crosstalk in erbium-doped fiber amplifiers," Optics Lett, vol. 14, pp. 880-882, 1989.

[14] E. L. Goldstein, V. da Silva, L. Eiskildsen, M. Andrejco,

and Y. Silberberg, "Suppression of dynamic cross- saturation in multiwavelength lightwave networks with

inhomogeneously broadened amplifiers," IEEE Photon.

Technoi. Letts., vol. 5, pp. 937-939, 1993.

REFERENCES

[1] M. J. Karol, G. Hill, C. Lin, and K. Nosu, eds.. Special

Issue on Broadband Optical Networks, IEEE J. Light- wave Technology, vol. 11, May 1993.

[2] C. R. Giles and E. Desurvire, "Modeling erbium-doped

fiber amplifiers," IEEE J. Lightwave Technology, vol. 9, pp. 271-283, 1991.

[3] E. L. Goldstein, L. Eskildsen, C. Lin, and R. E. Tench,

"Multiwavelength propagation in lightwave systems with

strongly inverted fiber amplifiers," IEEE Photon. Technoi. Lett, vol. 6, 1994.
OUT-OF-BAND ON-LINE MONITORING OF THE OPTICAL NETWORK

E. Cottino f*), M. Artiglia f**

j

(*) AET Telecomunicazioni - Via Sestriere 130, Rivoli - Cascine Vica, Torino (Italy)

(**) CSELT - Via G. Reiss Romoli 274, Torino (Italy)

Abstract

The state of the art ofmonitoring systemsfor the external optical plant is reviewed.

The various approaches proposed in the literature are discussed emphasizing the role of

in-service non-intrusive monitoring systems, whereby the important function of

preventive maintenance is accomplished. Existing systems and equipment are described

with particular emphasis on the first out-of-band on-line monitoring system ever realised

developed in Italy

1. Introduction

A substantial improvement in the quality of telecommunication services is expected

because of the superior performance of fibre optic cables. However, the huge amount of

information that can be carried by systems based on optical fibre is setting requirements on system reliability which are much more stringent than usually encountered in

traditional systems based on copper wire, particularly in view of the introduction of new

broadband services. In today's optical telecommunication systems, surveillance is made only on the

received signal by measuring the Bit Error Ratio of the transmission, according to ITU-T

(former CCITT) recommendations. No supervision of the physical carrier is considered,

contrary to what happens with copper cables. As a consequence, repair is time

consuming, costly and involves long service interruptions. Such situation, which can

hardly be tolerated in inter-office trunks maintenance, may become intolerable for

subscriber loop maintenance, where typically a very large number of fibres is involved

and maintenance made while the system is in-service is preferable.

Supervision of the optical carrier offers a very promising solution to these

problems [1-7], allowing to recognise immediately whether the fault is to be ascribed to the optical carrier and permitting the supervisor to localise promptly the damage with an

efficient exploitation of field work. Another important advantage is that by a suitable

choice of the maintenance wavelength window a continuous monitoring of the

performance of the installed fibres is possible, which enables the supervisor to prevent

degradation of the optical properties of the link and repair them before faults occur. Scope of the paper is to review the state of the art of in-service non-intrusive

monitoring techniques, with particular emphasis on an original technique developed in

Italy, called VDO (Valutazione del Degrado Ottico , Testing of the Optical Degradation of

the fibre plant). Advantages and drawbacks of the various techniques will be discussed;

results of the first field trials will also be presented and commented.

2. Basic principles

The primary scope of remote monitoring of the external plant is that of detecting

quickly from the central office a trouble or a critical situation which could lead to a fault,

thus reducing the time in which the system is out of order. The general structure of an

external plant maintenance system is comprised of three main parts: the testing system,

the diagnostic system and the data base system. The testing system should be able to access automatically the fibres of a given cable and remotely measure their characteristics.

The testing systems are distributed along the optical network. In principle, all the fibres

of a link should be taken under surveillance, without affecting service, for as much time

as possible in order to detect anomalies as soon as they occur. The testing equipment may

be for example an OTDR machine or a power meter plus suitable optical sources for

measuring fibre attenuation and/or other type of equipment (such as sensors for strain,

temperature or humidity). Then, the detected anomalies are signalled to the diagnostic

system, placed in the office, which processes the information and automatically requires,

if necessary, further measurements on the damaged fibre(s) in order to understand nature

and location of the fault. In this way network operators are able to analyse promptly the

25
trouble and initiate maintenance and administrative activities to accomplish repair and

return to service in an efficient manner. Besides these corrective tasks, another important

possibility opened by in-service surveillance systems is that of preventive maintenance.

By means of a real-time evaluation of the parameters of in-service and not-yet-in-service

links, early degradation of the optical properties of the fibre line can be prompdy detected

and removed, before the occurrence of service interruptions. This task can be

accomplished very well by monitoring the fibre in the 1600-1650 spectral region as discussed below [6]. Advantages of out-of band monitoring versus traditional methods

are summarised in Table 1. Then, information about faults is transmitted to a storage

centre so as to create a data base containing all data pertaining to the network

configuration. Automated and prompt input of this extremely voluminous and constantly

changing data is mandatory. Manual update of the data base, besides requiring an enormous amount of manpower, is much risky, because of possible errors and delays

which may cause the data base fail to reflect the actual status of the system. Finally the

system must fit into the Telecommunication Management Network (TMN) and thus

comply to existing TMN recommendations and standards.

The actual configuration and philosophy of the monitoring system will depend on many factors, such as the network characteristics and architecture, transmission

equipment and operation windows.

3. Proposed approaches to remote optical plant monitoring

Various approaches to remote monitoring of the optical network have been

proposed in the literature [2-6]. In a first approach, the aim is to detect only macroscopic damages of the optical

cable, which could impair the proper operation of a large number of lines, as shown in

Fig. la. To this end, the monitoring system make use of one or more fibres of the cable

(one fibre per cable groove, for example) dedicated to the transmission of the monitor

signal (dark-fibre approach). These systems cannot detect troubles of in-service fibres

supporting the traffic, but only catastrophic damages of the cable itself. A drawback of

the method is that a fraction of the fibres of the cable cannot be used for data

transmission, thus reducing the transmission capacity of the cable itself If the whole

cable is to be put in-service, monitoring becomes impossible.

To overcome the limitations of the dark-fibre approach, different approaches have

been devised which make use of in-service fibres to transport the monitor signal (see

Fig. lb). To avoid disturbing the transmission, the monitor signal must propagate in die

fibre at a wavelength different from that used for the traffic. Therefore, suitable

Wavelength Division Multiplexers (WDM) are to be used for monitor signal injection into

each fibre. At the end of the line the monitor signal is demultiplexed and detected. In this

way all fibres in the cable are kept under control. In this approach, the crucial element is the WDM, which must feature low crosstalk, low loss and low cost. As far as the

allocation of the monitoring wavelength is concerned, the monitor signal can be

transmitted at 1310 if the cable only supports systems operating at 1550 nm and

viceversa if the supported systems operate in the second window. Problems may occur when the the cable supports mixed systems; in this case, information about the

transmission wavelength on a given fibre should be available to the monitoring system,

which should implement two optical sources and a double set of WDMs for switching on

the right wavelength. Further problems may occur for WDM systems where, for

example, telephony is transmitted at 1310 nm and CATV is distributed at 1550 nm.

Allocation of a monitoring window outside the two spectral regions commonly used for

transmission would solve all these problems, allowing the monitoring of full cables

irrespective of the system characteristics.

Various transmission and detection techniques can be used for the monitor signal.

A first class of approaches use point-to point transmission, with transmitter separated

from the receiver (Fig. 1 .c). The second is based on Optical Time Domain Reflectometry

(OTDR) which allows for one-end access to the fibre (Fig. Id). In this case, very low

crosstalk WDMs are mandatory, since OTDR pulses are strong and may disturb the

transmission.

The use of point-to-point transmission of the monitor signal makes it possible to have a continuous monitoring of the global attenuation of the fibre link under test. The

26
measured attenuation, may, for example, be compared with some stored reference values

corresponding to proper system operation, so as to detect line interruptions or

degradation of attenuation characteristics with respect to reference. This method is well

suited for the maintenance of trunk network links, where the transmitter and the receiver

are located in the exchange offices at the ends of the line to be monitored. To reduce the

costs, the same receiver and/or transmitter can be shared by many fibres by using, for

example, a suitable mechanical optical switch. In this way the fibres of a cable can be

monitored cyclically by means of a fully automated equipment. As far as the distribution

network is concerned, this method is practical only in a ring architecture; in fact, the

dispersion of the many remote terminals in tree-structures distribution networks (such as PONs) makes the implementation of point-to-point monitoring quite difficult. On the

other hand, in a tree architecture, surveillance by monitoring a dark-fibre may be applied

to the feeder cable, from the office to the first distribution point.

Monitoring of the fibre line by means of OTDR equipment has the advantage of

providing information about possible reflections and behaviour of the attenuation along

the line, thus enabling possible critical points to be localised. Moreover, OTDR

equipment requires access to only one fibre end and therefore may be used for

monitoring tree-structured networks, even though special care must be exercised in

analysing the OTDR trace when passive branching points are present. On the other side,

the equipment is rather complex and costiy. An effective way of reducing costs is using a mechanical switch to select periodically the fibre to be monitored; this allows a single

OTDR equipment to be shared by large number of fibres. Some OTDR issues are still under study in view of their application in monitoring systems, such as improvement of

spatial resolution, for subscriber loop applications, extension of dynamically range, for

long distance, faster speed of operation and availability of optical sources if out-of-band

surveillance windows are allocated.

Also the possibility of monitoring installed fibre stress by means of Brillouin

Optical Time Domain Analysis has been proposed and is under study [7].

4. Equipment and field trials

The first generation of remote fibre test systems is already on the market, having

been introduced in the early 90's by fibre optic test and measurement equipment

suppliers. This equipment are likely to be deployed first in inter-office and feeder plant

and only later on in high-density fibre areas such as the subscriber loop. Telephone

network operators in USA, Europe and Japan are either deploying these machines or conducting field trials. The leading country today is Japan, where since 1990 nearly 300

monitoring systems have been installed, mostiy for long-haul trunk routes [1 1]. The OTDR technique is the one most frequentiy used in the surveillance systems

already developed or under study. This is the case of AURORA (Automatic Optical Fibre

Operations and Support System) realised by NTT [8,9], the first optical plant

maintenance system ever developed and deployed. AURORA provides automatic

continuous monitoring of fibre attenuation of in-service fibres from the value of suitable

reflections from the line terminals. AURORA, akeady deployed in Japan, is operating at 1310 nm for long-haul systems, where third window fibres are installed, and at 1550 nm

in the subscriber loop, where distribution is performed in the second window.

A general approach similar to this is applied in equipments developed by USA

suppliers, using dark or in-service fibres, although at a lower level of complexity.

Interesting advantages are offered by the OTDR based surveillance system, called

VDO, developed in Italy by AET [12]. VDO has been devised to fit the Operation and

Maintenance System of the Italian telephone operating company. In particular VDO is able to detect variations of the attenuation of fibre, connectors and splices which could

affect the service. The main structure of the system is shown in Fig. 2. As can be seen

the system is modular in order to gain flexibility and fit already existing office

infrastructures. The main feature of VDO is that it is the first surveillance system which

allocates an out-of-band monitoring window, placed in the long wavelength spectral

region beyond 1600 nm. The reason for this choice is twofold. In the first place, the

Italian optical network is based on two types of optical fibre, standard dispersion

unshifted fibres operating around 1310 nm and dispersion shifted fibres operating around

1550 nm. Secondly, in the spectral region around 1600 nm fibres are more sensitive to

27
bending and stresses, so anomalies in transmission properties can be more easily

detected. A wavelength of 1625 nm has been chosen, in order to measure fibre

attenuation values very close to those around 1550 nm and guarantee a good isolation at the data transmission wavelength. To support these statements, a comparison of

attenuation and splice loss measurements at 1550 nm and 1625 over a wide sample of

installed optical fibres is given in Fig. 3 and 4. As can be seen, the differences are very

small.

The system features an originally developed 240 fibre high-stability mechanical

switch; an OTDR equipment implementing also a laser source at 1625 nm; WDMs able to multiplex or demux 1310 nm, 1550 nm and 1625 nm signals with very high isolation.

The VDO operates on in-service fibres and spare fibres and gives periodically the

behaviour of the total attenuation of the monitored link signalling occurring anomalies. A

number of VDO machines has been deployed in Italy on the long distance network and

field trials are in run. Transparent to data transmission has been demonstrated by field

measurement of the system BER. BER curves of a 2.5 Gbit/s system measured in the

field before and after the installation of the monitoring systems are given in Fig. 5,

showing no significant degradation of the quality of the transmission after deployment.

The possibility of preventive maintenance, thanks to monitoring at 1625 nm, has also

been demonstrated in practice. Figure 6 shows the OTDR traces of a fibre placed in an

installed cable which was accidentally damaged during operation. Trace 1) is the OTDR

reference trace at 1625 nm before damaging. Traces 2) and 3) are the OTDR traces at 1625 nm and 1550 nm respectively, taken after damage occurred. As can be seen the

1550 nm trace doesn't show any degradation, while the 1625 nm trace shows a sudden

drop of attenuation at the damage location. The attenuation drop is well inside the system

margin, so no service interruption occurred. However, the fibres of the cable have been

and still are subjected to an anomalous stress which could favour fibre breaks with

consequent risk of service interruptions. This situation could not be detected by

monitoring the cable at 1310 nm or 1550 nm, as AURORA or other proposed systems

do.

5. Conclusions

The widespread use of optical fibre in long distance and inter-office network and

the progressive introduction of optical fibre in the subscriber loop originated an ever

increasing interest in test facilities of the fibre plant which guarantee a continuous

monitoring of the fibre lines and operate when the system is live. Various solutions have

been proposed. The most promising approach seems to be the continuous monitoring of

in-service fibres by means of OTDR equipment operating in the 1600-1650 nm window.

In this way, non intrusive preventive maintenance of the optical plant is possible. The

first out-of-band in-service monitoring system, developed in Italy, has been presented.

This area is rather new and much work is still to be done. Also other fibre properties

besides attenuation, such as fibre strain, may be monitored, permitting on-line prediction

of fibre reliability, and, last but not least, proper standards for the massive introduction

of remote system monitoring in the telecom network are to be studied and issued.

RcfcrcncGS

[I] F.A. Huzarik et al, Proceedings of IW&CS pp. 158-162 (1988)

[2] H. Hornung et al.. Proceedings of ICC 90, p. 341.5 (1990)

[3] Y. Koyamada et al., Proceedings of ICC 90, p. 341.1.1 (1990)

[4] H. Koga et al., Proceedings of ICC 91, paper 11A.3 (1991)

[51 P. Keeble et al. Proceedings of ICC 9 1 , paper 1 1 A.2( 1 99 1

)

[6] I. Sankawa et al., Proceedings of IW&CS 92, pp. 81-87 (1992)

[7] T. Horiguchi et al., J. Lightwave Technol., vol. 7, pp.1 170-1 176 (1989)

[8] H. Matsumoto et al.. Proceedings of IW«&CS 91, pp.787-791 (1991)

[9] N. Tomita et al.. Proceedings of EFOC 91, London, pp 130-134 (1991)

[10] H. Takasugi et al.. Proceedings of IW&CS 90, pp.623-628 (1990)

[II] Marketing Intelligence, Fiberoptic, April 25, 1 194

[12] E. Cottino et al. , Proceedings of IW&CS 93, pp. 799-804 (1993)

28
a)

MAINTENANCE PURPOSE TRADITIONAL METHOD OUT-OF-BAND

MONITORINQ

p EVAJ.UATX3N OF M -SERVICE L»i< R PaR»J*TERS (FBflES. SPUCES. NO YES (from E COM^CTOflS) MANTENANCE CENTRE ) V

E Evaluation of not-n-sehvice YES YES (From N

T

LH( PARAMETERS {Measure»«nt n Loco) Mantehance Centre )

1VEarly Degrawtkx Detectkjn NO (May Be) YES

E Tfce.y Degradation Removal NO YES

C0 Detectcnof Damage Not Precbe YES

RR

(Man/Nature, Type, Location)

E Detectcn TilC HOURS MINUTE c

T After Re pa« Test A^o MANUAL AUTOMATIC

RECOfiO»« (From Maktenance >

Ul 1

Centre)

Table 1b)

Transmission Equipment

In-ser 'ic« fibres

Monitoring

System

Dark

fibre

Central Office

OpticaJ Cable

Transmission

Equipnr»4it ^

In-serN Ic* fibre! 71

WDM

Monitoring

System

Central Office

Optical Cable

d)

OTDR

Equipment

Office Fbre Cabte

Fig. 1 - Review of moniioring techniques. See Section 3.

TMO

OPTICAl. FILTER

TX

TRANSMISSION RACX

DCN

OS

P.O.

OPTICAL SWITCH

TMO

DATA

BASE PCI

ws

OPERATOR

OPTICAL FILTER

TX

EQUIPMENT

TOAHSMISSION RACK

Fig. 2 - Schematic of the VDO system. PCT: local system controller; TMO: Optical Monitoring Rack Components; DCN: Data Commimication Network;

MD: Mediation Device.

29
«>

M o « o « n o nucomo

'SEouo

V3Q 3 >

•a

4> 602 O

^ir I o b "5 S

ir: o -Si « c o i3 ^ « o S S ^

•a c3 vj C CO

4> {« « W U _ C3 Q •3 p

.O s >o \o

MEc

T3OCo

CO iEoo

CO

r ?' lU8 _ k 2Q

on

JOO co

^. ^

V)

a 2J o 2

31VU Hoa«3 Jja

30
Multi-100 Gbit/s optical waveform measurement

using sum-frequency-generation optical sampling

with subpicosecond supercontinuum pulses

H. Takara, S. Kawanishi, T. Morioka, K. Mori and M, Saruwatari

NTT Transmission Systems Laboratories

1-2356 Take, Yokosuka, Kanagawa, 238-03 Japan

Introduction

Ultrahigh-speed optical waveform measurement becomes the key to evaluating ultrahighspeedAong-distance optical fiber transmission characteristics [1]. Optical sampling based on

optical nonlinearity is expected to realize the waveform measurement with subpicosecond

temporal resolution. To date, several optical sampling methods have been reported based on

sum-frequency-generation (SFG) [2] [3], nonlinear optical loop mirror (cross-phasemodulation) [4] and four-wave mixing [5]. However, the temporal resolution was limited to 6~150 ps by the sampling pulse sources such as mode-locked external cavity laser-diodes,

gain-switched laser-diodes and Nd:YAG lasers. Recently, we reported 0.63 ps resolution 100

Gbit/s optical waveform measurement using SFG optical sampling with subpicosecond (0.55

ps) sampling pulses by supercontinuum (SC) generated in an optical fiber [6]. In this report, we demonstrate the successful waveform measurement of a 400 Gbit/s

optical signal with the improved resolution of 0.44 ps by shortening the pulse duration of the

SC sampling pulses down to 0.30 ps. This optical sampling system can measure, with good

linearity, optical signal waveforms over wavelength ranges of more than 25 nm around 1.55

|im. Also discussed is the S/N characteristics of the proposed method.

Experiments

Figure 1 shows the experimental setup, which is based on that of the previous report [6].

Summarizing the configuration of the previous setup, 3.0 ps, 0.9 nm bandwidth optical pulses

from a mode-locked Er-doped fiber (ML-EDF) laser were converted into SC pulses in an SC

fiber [7]. 1559 nm-wavelength subpicosecond sampling pulses were obtained by filtering the

SC pulses. The signal pulses were also generated by the same SC technique. The repetition

rates of the signal and sampling pulses were set to 6.3 GHz and 6.3 GHz minus the frequency

offset of 50 Hz, respectively. The 6.3 Gbit/s signal was multiplexed to the 100 Gbit/s signal

with a planar-lightwave-circuit multiplexer (PLC-MUX) [8]. The samphng and signal pulses were combined by a polarization beam splitter (PBS) and injected into a KTiOP04 (KTP)

crystal. Type II phase-matching SFG was used to eliminate the unwanted second harmonic

light generated from the sampling or signal light because two orthogonal polarization lights are needed for phase-matching [3]. The crosscorrelation signals (the SF light) at 777 nm were

detected with a photomultiplier tube and the sampled waveform at a repetition frequency of 50

Hz was displayed on an oscilloscope. The details are described in the previous report.

31
To upgrade this basic setup, we shortened the sampling pulse duration to improve the

temporal resolution; the goal was measurement at 400 Gbit/s. Shorter sampling pulses were

obtained by compensating the chirping of the SC pulses with a 190 m, 1.3 |im zero-dispersion

fiber (Comp. fiber) and by filtering these pulses with a 3nm Lorentzian shape bandpass filter (BPF) [9]. An autocorrelation trace and the optical spectrum of the sampling pulses are shown

in Fig. 2. The calibrated pulse duration tsam and spectral bandwidth were 0.30 ps and 3.1 nm,

respectively. The time-bandwidth product At-Av was calculated to be 0.12, close to that of the

single sided exponential decay shape transform-limited pulses (At-Av = 0.11). The peak

power of the sampling pulse Psam was amplified to 1.06 W by an Er-doped fiber amplifier

(EDFA).

In order to generate the 400 Gbit/s optical signal from the 100 Gbit/s optical signal, two

polarization maintaining fibers (PMl and PM2) with polarization dispersions of 15 ps and 7.5

ps, respectively, were used as an additional 4:1 multiplexer. At the connecting points (A, B

and C in Fig. 1), the principal axis offsets were set to 45 deg. in order to make the single

polarization state 400 Gbit/s signal. The original signal pulse duration tsig of 0.45 ps was

increased to 1.4 ps by the 6.3 Gbit/s to 400 Gbit/s multiplexing. This is mainly attributed to the influence of filtering with a 10 nm BPF to eliminate the amplified spontaneous emission of

EDFA at the polarizer output. The amplified peak power Psig was 40 mW.

Results and discussion

Figures 3 shows the measured waveform of the 400 Gbit/s optical signal. This optical

sampling system well measured the data pattern "0100101 101 101 10" with the measured pulse

duration t^eas of 1-5 ps. The temporal resolution t^es of the measurement system was given by,

^res ~ itmeas ^ " hig ^ 'jsig ^ )^^^ (1)

where jsig is the timing jitter of the signal pulses. From the measured jsig of 0.31 ps, tres was

estimated to be 0.44 ps. This value well agrees with the calculated value considering the pulse

duration {tsam = 0-30 ps) and the timing jitter (Jsam = 0-30 ps) of the sampling pulse (0.42 ps =

iitsam )2 + (/"W2)l/2).

As seen in Fig. 3, background-free detection was realized because of type II phasematching. Please note that no optical filter was used for extracting the SF light in this

experiment. This means that this method can measure the optical signal waveform at any

wavelength, even the same wavelength as that of the sampling light. The signal wavelength

dependency of the SF light level is shown in Fig. 4. The change of the SF light level was less

than ±5 % for the wavelength range from 1535 nm to 1560 nm. This result shows that this

optical sampling method has a wide signal wavelength tolerance of more than 25 nm around

1.55 |J.m. Fig. 5 shows the relationship between the crosscorrelation signal (SF light) levels

and the input signal peak powers. The crosscorrelation signal levels were proportional to the

signal peak power up to 400 mW. It is clear that this scheme realizes optical waveform

measurement with good linearity.

The SN ratio of type II phase-matching SFG optical sampling is given by [3]

32
^

- ^ 'HsfgPsamPsig

where r\sfg is the SFG conversion efficiency, K is the averaging number, 77^ is the quantum

efficiency, A is the wavelength of the SF Hght, h is Plank's constant, c is the velocity of hght,

B is the bandwidth of the photomultiplier, and ris the pulse duration of the SF light. Here,

we assume that shot noise is much larger than any other noise, such as thermal noise in the

photodetector. r]sfg of the KTP crystal was measured to be 6 x 1 0"^ W" ^ . X ,B and t were

777 nm, 1 kHz and 0.30 ps, respectively. The calculated SN ratio of one shot {K = 1) is 0.3

cind this value is too small to measure optical signal waveforms. In this experiment, the SN

ratio was improved by the averaging performed in the photomultiplier. Accordingly, the

averaging number K was the ratio between the SF light repetition frequency frep of 6.3 GHz

and the photomultiplier bandwidth B of 1 kHz {K =frep /B = 6x 10^), so the high SN ratio of

2 X 10^ was realized.

Conclusion

We have demonstrated the successful measurement of a 400 Gbit/s optical signal waveform

by using type II phase-matching SFG optical sampling with SC pulses. The temporal

resolution was improved to 0.44 ps by utilizing 0.30 ps chirp-compensated SC sampling

pulses. This optical sampling system can measure, with good linearity, optical signal

waveforms over wavelength ranges of more than 25 nm around 1.55 |a.m.

Acknowledgment

The authors would like to thank Dr. Hideki Ishio for his encouragement.

References

[1] S. Kawanishi, T. Morioka, O. Kamatani, H. Takara and M. Saruwatari, Tech. Dig.

OFC'94, PD-23 (1994)

[2]. T. Kanada and D. L. Franzen, Opt. Lett., 11, pp. 4-6 (1986)

[3]. H. Takara, S. Kawanishi, Y. Yamabayashi and M. Saruwatari, Transactions of lEICE

(B-1), J75-B-1, pp.372-380 (1992) (in Japanese)

[4]. B. P. Nelson and N. J. Doran, Electron. Lett., 27, pp.204-205 (1991)

[5]. P. A. Andrekson, Electron. Lett., 27, pp. 1440-1441 (1991)

[6].H. Takara, S. Kawanishi, T. Morioka, K. Mori and M. Saruwatari, CLEO'94, CPDIO

(1994)

[7].T. Morioka, S. Kawanishi, K. Mori and M. Saruwatari, OFC'94, PD-21 (1994)

[8].K. Uchiyama, H. Takara, S. Kawanishi, T. Morioka, M. Saruwatari and T. Kitoh,

Electron. Lett., 29, pp. 1870- 1871 (1993)

[9].T. Morioka, S. Kawanishi, K. Mori and M. Saruwatari, submitted to Electron. Lett.

33
6.3 GHz

^ , BPF

SC Comp.--^

fiber fiber

signal source

1550 nm jn6 PM2

PLC

MUX

m X 16 >*V J

B C 10 nm

mixer

sampling pulse source 6.3 GHz - 50 Hz ' 0.3 pS

6,3 GHz - 50 Hz

J^r-. SC fiber Comp. o

1559 nm : bpf]—|>-4-

400 Gbit/s

1 .4 ps

»r\ipBS

nm

50 Hz

oscillo- scope

photo

multiplier

777 nm

KTP

crystal PC

Fig. 1 Configuration for SFG optical sampling

0.30 ps

(calibrated)

Delay (ps)

Fig. 2 Optical spectrum and autocorrelation

trace of the sampling pulses

010010110110110

Fig. 3 Sampled waveform of

the 400 Gbit/s optical signal

1530

Fig. 4

1540 1550

Signal wavelength (nm)

1560

Signal wavelength dependency of

the SF light intensity

-5- >« % 3 c

cZ 2

O)

0)

1 1 1

ox

^ig = 1535 nm

"kaam = 1559 nm

1

1

Fig. 5

100 200 300 400

Signal peak power(mW)

Relationship between intensity of

the SF light and signal peak power

34
Characterizing Modal Interference in

Field Installable Single-Mode Fiber Connectors Incorporating Short Fiber Stubs

D. O. Harris and R. A. Throckmorton

SIECOR CORPORATION

489 Siecor Park - HE, Hickory, NC, 28603-0489

1. Introduction

In many fiber-optic applications, elimination

of excess cable lengths or emergency repairs necessitate installation of connectors onto deployed optical fibers. Because conventional connectors designed for factory installation must be cured at high temperatures and meticulously polished, they are not practical for field installation. As an alternative, field installable fiber- optic connectors have been developed that consist of a

short (< 2 cm) fiber stub preinstalled into a connector housing [I]; the connector is installed in the field by

mechanically splicing the stub to a deployed fiber. The proximity of single-mode fiber joints in such connectors can produce modal interference which may

cause a significant dependence of static connection loss on both source coherence and operating wavelength.

In order to understand the field operation of

these connectors and subsequently establish accurate test procedures indicative of their field performance, we must

first understand the mechanisms which cause modal

interference. The theory conventionally used to describe modal interference in single-mode fiber joint pairs was

developed to explain modal noise in applications where

the operating wavelength is slightly below cutoff and

the distance between joints is greater than 0.5 meters

[2]-[4]; in this case, initially orthogonal LPn modes

generated at the first joint may couple within the intermediate fiber so that the energy they carry is redistributed upon arrival at the secondjoint. This same

theory has since been used for applications where the distance between joints is of the order of centimeters

[5]-[7]. In the latter case, modal interference can affect the loss of the joint paii; but no modal noise is present

[5], [6]; moreover, there is minimal mode coupling or polarization conversion within the short intermediate stub fibei; so modes which are orthogonal just beyond

the first joint tend to maintain their independence

throughout the entire length of the stub. Given these subtle differences, we have modified the theory for the aggregate transmission efficiency of a closely spaced

single-mode fiber joint pair operating slightly below

cutoff This revised theory shows that the LPq i to LPii dominated modal interference is dependenton the azimuthal angle between the lateral offsets at the two

joints. The azimuthal effect is verified experimentally

input output

fiber fiber stub fiber

(a)

<P

(b)

Figure 1. (a) Longitudinal and (b) transverse cross section

of a two-joint single-mode fiber system wbiicli generates

modal interference. For a field installable connector

incorporating a short fiber stub, the joint separation is on

the order of a centimeter.

not only for operation below LPj j cutoff wavelength,

but also at wavelengths above cutoff where interference occurs between the LPq j and radiation modes [7]. From our revised theory, we derive an approximate formula for the "modulation index," which

expresses the uncertainty of connector insertion loss due

to modal interference as a percentage of the expected

loss when no modal interference is present [8]. We

further implement a method for measuring this parameter, and use results of such measurements to characterize modal interference for a sample population

of stub connector assembUes.

2. Transmission Efficiency

We analyze the case of two closely spaced

single-mode fiber joints with lateral misalignment as shown in Figure 1 for the condition where the operating

wavelength is below cutoff. Using the Gaussian

approximation for both LPqj and LPj j mode ampUtudes

[2], [9], overlap integrals yield the following analytical

results for couphng efficiencies at the first joint:

35
LPqi —> LPoj:

LPoi

LPoi

LP„:

higher order modes:

^1

-rji In

(1) The LPqi to LP^ coupling efficiency given in equation (1) agrees with previously published empirical

results [10] to better than 1% when T]^ > 0.8 (< 1 cB

single-modejoint loss, which encompasses the range of

practical interest). Equation (1) also indicates that for

7]i > 0.8, the LPji mode power just beyond the first joint is at least an order of magnitude greater than the

total power contained in the higher order modes; this suggests, as does conventional theory, that intermodal

interference beyond the secondjoint will be dominated

by interaction between the LPq i and LPj i modes of the stub fiberGiven the above result, we consider only

coupling between the LPqi and LP^ modes to approximate the aggregate transmission efficiency of a

closely spaced, low-loss joint pair at wavelengths below

cutoff. We first calculate the normalized LPq j mode

amplitude beyond the second joint through analytical

evaluation of correlation integrals between modes at tlie joints, while also accounting for mode attenuation and

optical path differences in the stub fiber; for this calculation, we assume that no mode coupling or polarization conversion occurs within the stub. The

square modulus of our normalized amplitude then yields

the following expression for LPqi transmission

efficiency through the low-loss joint pair:

T/= T]iri2\l + ln r7j In r/jlO 10 cos^ (p

-2r{Anl)^\n 7]^ In 10 ^° cosipcos

InAnl

(2) where and r]2 are the LPqi to LPqi efficiencies at the

first and second joints, respectively, is the LPn

mode attenuation in the stub fiber, / is the length of the stub fiber, cp is the azimuthal angle between the lateral offsets at the fiist and secondjoints , r{AnD is a source autocorrelation for the optical path length difference between the LPqi and LPi 1 modes in the stub which

ranges from 0 for incoherent illumination to 1 for completely coherent illumination. An is the refractive index difference between the LPqi and LPj 1 modes in the stub fiber, and A is the free-space operating

wavelength.

As stated in the introduction, a significant difference between this equation and the theory

white source chofper monochromator

5m

3-axis

piezoelectric drives

matched clad SMF

17mm

matched clad SMF stub

5m

matched clad SMF

detector

lock-in amplifier

Figure 2. Apparatus for demonstrating the azimuthal

dependence of modal interference.

convenfionally applied to modal interference is the dependence of the second and third terms on the azimuthal angle ((r, this dependence arises from the assumption that polarization alignment and mode

structure are maintained in the stub, which is the expected behavior for a short stub with no bends.

To verify the azimuthal dependence, we used

the apparatus in Figure 2 to measure spectral loss for one joint pair in the region below the LP^ cutoff wavelength of its the stub and for anotherjoint pair in the spectral region above stub cutoff; both joint pairs are index matched and separated by approximately

17 mm of matched cladding fiber The cutoff wavelength for the stub fiber of the first joint pair as measuredby the EIA method [11] is 1289 nm, while the EIA cutoff for the second stub is 1209 nm; the respective cutoff wavelengths of these stub fibers in

their deployed condition are estimated to be 1435 nm

and 1295 nm as determined by the interference method

described in Ref. 12. For both pairs, we introduce approximately 0.2 dB of loss due to lateral offset at each

joint for the three cases where (p = 0°, 90°, and 180°. Spectral loss from 1260 nm to 1360 nm is shown in Figure 3(a) for the first joint pair; as equation

(2) predicts for operation below cutoff wavelength,

spectral loss for the two parallel offset cases ((jp = 0° and

(p = 180°) have approximately the same high levels of

modal interference but nearly opposite polarity, while

modal interference practically vanishes when the offsets are orthogonal {cp = 90°). Figure 3(b) contains a plot of

spectral loss over the 1500 nm to 1600 nm region for the second joint pair, representing operation above

cutoff wavelength. In this case, interference occurs 36
1260 1280 1300 1320 wavelength (nm)

(a)

1340 1360

es3

i

e

s

1500 1520 1540 1560 1580 1600 wavelength (nm)

(b)

Figure 3. Measured spectral loss (a) below and (b) above

cutoff wavelength for closely spaced, single-mode fiber joint pairs with approximately 0.2 dB of lateral offset induced loss at each joint and offset azimuthal angles of 0°, 90°, and 180°.

between the LPqj and radiation modes, but the azimuthal effect still exists; this is consistent with the notion that lateral offset causes light to be coupled

primarily to radiation modes having amplitudes with

odd synmietry about the fiber axis.

3. Modulation Index

When individual joints are reasonably efficient

(7]i, TJ2 > 0.8), the product rjj in equation (2) is very nearly the aggregate transmission efficiency of the joint

pair under incoherent illumination. If we designate

TjjT]2 as T). , then we may also write 7]j = (t]-„^)^ and =

(^,„P^

'^^ where ^ ranges from 0 to 1. Further,

Li = gL-^^, andLj = (l-,?)^/„c' where andL2 ate the losses at the first and secondjoints, respectively,

andL-^^ is the approximate aggregate incoherent loss of

the joint pair (i.e., loss under incoherent illumination).

Hence, the parameter g represents the incoherent loss allocation between the two joints.

The extreme values of aggregate insertion loss over the spectral range of interest can be represented as ^max = Li^^(l+m) andL„„„ « L,„c(l-m) [8], where

m is called the "modulation index;" accordingly.

(dB)

(3)

is the approximate incoherent loss for the spectral range of interest and

m -

(4)

For a low-loss single-mode joint pair (L,„^ < 1 dB), we

can derive the following approximation for modulation

index based on equation (2):

»i = 2r(Zl«/)-^g(l-g)10 20 coscp

(5)

From this expression, we see that modulation

index is dependent on the source coherence, loss allocation between the joints, LPi 1 mode loss in the stub, and the azimuthal angle between the offsets. Since the modulation index represents the loss uncertainty from modal interference relative to incoherent loss and is, in theory, independent of

incoherent losses < 1 dB, it is a good candidate for characterizing modal interference. To demonstrate the application of modulation

index in characterizing modal interference, we fabricated 28 field installable connectors having 17 mm matched

cladding stubs with EIA cutoff wavelengths of

1203 nm; the estimated cutoff wavelength for these stubs in their deployed state is near 1300 nm. Spectral

loss was measured by the cutback method for matings

of each of these connectors to a factory installed connector using the transmitter and receiver configuration shovm in Figure 2. Modulation indices for the 1260 nm to 1360 nm spectral range were calculated for the 28

connectors using equation (4); Figure 4 contains a

histogram of the modulation indices for these sample

connectors. Since modulation index is basically

independent of incoherent loss, histograms such as this can be can be used to describe connector insertion loss uncertainty due to modal interference in much the same way histograms of connector insertion loss are presently

used. For the current example, the average modulation

37
0.0 0.2 0.4 0.6 0.8 1.0 modulation index

Figure 4. Histogram of modulation index over the 1260 nm

to 1360 nm spectral range for 28 stub connectors. Since

modulation index is essentially independent of incoherent

loss, its distribution and sample statistics can be used to characterize modal interference in much the same way

insertion loss distributions are currently used.

index is approximately 0.18 while the maximum is 0.60; therefore, we would assert that the coherent

insertion loss uncertainty due to modal interference for connectors from this population will be about ±18% on

average, with a worst-case of ±60%.

4. Conclusion

We have have provided a theoretical basis and

demonstrated that modal interference in a closely spaced

single-mode fiber joint pair is dependent on the azimuthal angle between the lateral offsets at the two

joints — an effect which is not predicted by

conventional theory. Using our revised theory, we

deduce a parameter called the modulation index that can be used to characterize modal interference. A

measurement technique for determining the modulation

index of stub connectors is implemented, and the modulation index distribution obtained from

measurements on a set of connectors is used to characterize modal interference for the set. The authors wish to acknowledge technical discussions with C. K. Eoll, W. S. Jackman,

C. Saravanos, andO. I. Szentesi, as well as assistance with experiments from S. C. Grant and her staff.

References

[1] M. deJong, "Cleave and crimp fiberoptic connector

for field installation," Tech. Digest Conf. on Optical

Fiber Commun., paper ThAl, p. 139, 1990.

[2] S. Heckmann, "Modal noise in single-mode fibres operated slightly above cutoff," Electron. Lett., vol.

17, pp. 499-500, 1981.

[3] D. G. Duff, F. T. Stone, and J. Wu, "Measurements of

modal noise in single-mode lightwave systems,"

Tech. Dig. Conf. on Optical Fiber Commun., paper TuOl.pp. 52-53, 1985.

[4] M. Brandtner, D. Schicketanz, and C. Eoll, "Modal

noise due to second mode interference in singlemode

fiber systems," Proc. 36th IWCS, pp.622-628, 1987.

[5] G. A. Olson and R. M. Fortenberry, "Modal noise in single-mode fiber-optic systems with closely spaced

splices," Fiber and Integr. Opt., vol. 9, pp. 237-244,

1990.

[6] J. C. Goodwin and P. J. Vella, "Modal noise in short

fiber sections," J. Lightwave Technol., vol. 9, pp. 954-958, 1991.

[7] K. Abe, Y. Lacroix, L. Bonnell, and Z. Jakubczyk,

"Modal interference in a short fiber section: fiber length, splice loss, cutoff, and wavelength

dependences,"/. Lightwave Technol., vol. 10, pp. 401-406, 1992.

[8] W. W. Wood, unpublished work, 1993.

[9] A. W. Snyder and J. D. Love, Optical Waveguide

Theory, New York: Chapman and Hall, p. 307, 1983.

[10] L. Wei, R. S. Lowe, andC. Saravanos, "Practical upper

limits to cutoff wavelength for different single-mode

fiber designs," J. Lightwave Technol., vol. 5, pp. 1147-1155, 1987.

[11] Electronic Industries Association, "Cutoff wavelength

of uncabled single-mode fiber by transmitted power,"

EIA-455, FOTP-80, 1988.

[12] K. Abe, Y. Lacroix, Y. Cai, and L. Bonnell,

"Determination of LPj \ mode cutoff wavelength from

the modal interference pattern," Syinp. on Optical

Fiber Measurements, NIST Special Publication 839,

pp. 221-224, 1992.

38
Performance limits of coherent OTDR due to optical nonlinear effects

Hisashi Izumita, Yahei Koyamada, Shin-ichi Furukawa, and Izumi Sankawa

NTT Telecommunication Field Systems R&D Center,

Tokai, Ibaraki-ken 319-11, Japan. AbstractThis paper clarifies the limit of incident optical pulse power in coherent optical time-domain reflectometry

(C-OTDR) enhanced with optical fiber amplifiers. The critical pulse power, at which the performance of C-OTDR is degraded by the effect of optical nonlinear phenomena in a single-mode optical fiber, depends on the amplified optical

pulse waveform, the pulse width and the effective length of the fiber. This paper also demonstrates the high

performance of C-OTDR enhanced with EDFAs with 48 dB, 39 dB, and 29 dB single-way dynamic ranges after 2'^

integrations for pulse widths of 10 |is, 1 |is and 100 ns, respectively, Hmited by the effect of SPM or FWM.

1 . Introduction

Optical time-domain reflectometry (OTDR) is a commonly used technique for characterization and fault

location in optical fiber transmission systems. Many studies have been undertaken to extend the dynamic range of

OTDR [1], [2]. The use of erbium-doped fiber amplifiers (EDFAs) combined with a coherent detection technique is very effective in enhancing the performance of OTDR [3]-[6]. EDFAs have the potential for high gain and high

saturation output power. An optical pulse can be amplified to more than 1 W by using EDFAs. However, a high-power optical pulse with the narrow linewidth needed for coherent detection sfimulates optical nonlinear

phenomena in a single-mode optical fiber [7]. These phenomena significantly degrade the performance of coherent OTDR (C-OTDR). It is therefore important to clarify the critical power of these phenomena, which are the origin of

the single-way dynamic range (SWDR) limit of C-OTDR combined with EDFAs.

This paper describes theoretical and experimental studies on optical nonlinear phenomena stimulated by a high-power coherent pulse in a single-mode optical fiber, and clarifies the critical power for the degradation in performance of C-OTDR due to these phenomena. It shows the high performance of C-OTDR combined with

EDFAs for 10 jis, 1 |is, and 100 ns pulse widths.

2. Experiments

Experiments were performed to clarify the critical conditions of the incident optical pulse for degrading the

performance of C-OTDR due to optical nonlinear effects in the optical fiber.

2.1 Experimental set-up. Figure 1 shows the experimental set-up for C-OTDR with EDFAs. A DFB-LD was coupled to a 1-km fiber

to reduce the linewidth to less than 10 kHz. The continuous-wave (CW) light from the DFB-LD was divided into two paths by a3-dB fused-fiber directional coupler (FCl): a signal path and a local oscillator (LO) path. The signal

pulses were produced by an optical waveguide switch made of LiNbOj (EO-SW) operating in pulsed mode. The pulse

period was 2 ms. The signal pulses were amplified by two post-amplifiers (EDFAl and EDFA2) and an in-line

amplifier (EDFA3). An acousto-optic switch (AO-SW), operating in pulse mode synchronized with the EO-SW,

eliminated the amplified spontaneous emission (ASE) from EDFAl. The optical frequency of the signal pulses was

shifted by 120 MHz at the AO-SW. The amplified signal pulses were launched into a 100-km test fiber. The

Rayleigh backscattered signals from the test fiber were mixed with the LO by FC2 and detected with a

double-balanced p-i-n-FET receiver. 2.2 Experimental results. Figure 2 shows the C-OTDR traces measured for a 1 |is pulse width. In trace (a), the incident pulse power was

21 dBm. The total loss of the test fiber was found to be 22 dB from the C-OTDR trace, which agrees with the total

loss measured with the optical power meter. In trace (b), the incident pulse power was 24 dBm. Compared with trace 39
(a), the attenuation curve of the optical fiber on the C-OTDR trace gradually became distorted from 10 to 100 km.

The total measured loss of the optical fiber was 24 dB, which is 2 <fi larger than that of (a). Figure 3(1) shows incident optical pulse waveforms with a pulse width of 1 |is and an average pulse power

over the pulse width of (a) 21 dBm and (b) 24 dBm. The power of the optical pulse amplified by EDFAs in the

saturation regime gradually deaeased from the head of the pulse to its tail, due to a reduction in the inversion

population in the erbium-doped fiber. The power gradient within the pulse width increased in proportion to the pulse

power. In these experiments, the power gradient within the pulse width was approximately constant, and the power

gradient k with the pulse width of x was given by

where p(t) is the optical pulse power at time t. From Fig. 3(1), the measured power gradient k for (a) was 2.9x10^* W/s and for (b) was 6.6x10" W/s.

Figure 3(2) shows the spectra of the beat signal between the transmitted optical pulse and the LO. These

spectra were measured with a conventional spectrum analyzer. In (a), the optical spectrum was in the form of the square of a SINC function with a center frequency of 120.4 MHz and a period of 1 MHz. The center optical frequency was shifted by 0.4 MHz in comparison with the optical frequency shift of 120 MHz at the AO-SW. In (b), the center

optical frequency was shifted by 1 MHz.

Figure 3(3) shows the optical spectra in the wavelength domain of the fransmitted optical pulse. In (a), the

center wavelength of the transmitted pulse was 1.5513 [im. No Stokes or anti-Stokes light was generated. In (b), two side-bands were generated, one on each side of the center wavelength, as a result of four-wave mixing (FWM).

These experiments show that the critical condition for degrading the performance of C-OTDR combined with

EDFAs is given by the optical frequency shift of the Rayleigh backscattered signals caused by the effect of SPM

(discussed in next section). The critical power gradient of the incident pulse amplified using the EDFAs was 2.9x10"

W/s and the pulse power for this critical condition was lower than the critical power for the other nonlinear

phenomena.

Experiments were also performed for a 100 ns pulse width. Figure 4 shows the C-OTDR traces. In trace (a),

the incident pulse power was 24 dBm. The total loss of the test fiber was 22 dB, which agrees with the experimental

results shown in Fig. 2(a). In trace (b), the incident pulse power was 25 dBm. The C-OTDR frace was distorted and

at 20 km it was shifted down by about 2 dB. The total measured loss of the test fiber was 2.5 dB greater than that of

(a). Figure 5(1) shows the incident optical pulse waveforms. The power gradient of the pulse was negligible for both (a) and (b). Figure 5(2) shows the specfra of the beat signal between the fransmitted signal and the LO. They

had the form of the square of a SINC function with a center frequency of 120 MHz and a period of 10 MHz. The

optical frequency shift of the transmitted optical pulse was also negligible in both cases. Figure 5(3) shows the optical specfra in the optical wavelength domain. In (a), the side-bands were very small.

In (b). Stokes and anti-Stokes lights was generated on either side of the center wavelength as a result of FWM. This

caused the incident optical pulse to interact parametrically with the amplified spontaneous emission (ASE). The

power levels of the Stokes and anti-Stokes light were 1 5 dB higher than the ASE level. The optical wavelength shift

(the optical frequency shift) between the signal and the Stokes or anti-Stokes light was about 0.5 nm (62 GHz) and

the Stokes bandwidth was about 0.25 nm (32 GHz).

The test fiber consisted of several 5-20 km dispersion-shifted single-mode fibers with typical dispersion D of

-1.50 ps/(km nm). From Eq. (10.3.10) in Ref. [8], the calculated frequency shift and the Stokes bandwidth at the

incident pulse power of 316 mW (25 dBm) were 147 GHz and 74 GHz, respectively. The reasons for the difference

40
between the measured value and the calculated one were some fluctuation of the dispersion in the test fiber and the parametric gain suppression of FWM caused by polarization mismatch between the signal pulse and the ASE.

These experiments show that the critical condition for a pulse width of 100 ns is given by FWM, which leads

to the generation of the Stokes and anti-Stokes light. The FWM critical power for the incident pulse amplified with

the EDFAs was lower than the critical power for the other nonlinear phenomena.

3. Discussion

When an optical pulse with a power gradient within the pulse width is incident to the optical fiber, the

optical frequency of the transmitted optical pulse passing through the optical fiber is shifted by self-phase

modulation (SPM) [9]. The Rayleigh backscattered signals have the same optical frequency shift as the optical pulse

propagating in the optical fiber. Since a stable frequency difference between the Rayleigh backscattered signal and the

local oscillator is needed for coherent detection, the optical frequency shift degrades the sensitivity of C-OTDR

because the center frequency of the signal moves outside the receiver band. When the critical frequency shift 5v, is defined to make the sensitivity degradation less than 0.05 dB, 6v^/B is 0.4 [6], thus the critical power gradient of the incident optical pulse is [6] ^1 =^5v,= -^0.4B 3t ^ n^Leff = njLeff , (2)

where _ l-exp(-aL)

^eff- G ' A is the effective cross section area, B is the receiver bandwidth, p(t) is the optical pulse power at time t, n2 is the nonlinear refractive index of the optical fiber, X is the wavelength of the incident optical pulse, L^ff is the effective

length, L is the length of the optical fiber, and a is its attenuation coefficient. Equation (2) shows that the critical power gradient is proportional to the receiver bandwidth and the inverse of the effective length. For a receiver bandwidth B of 1 MHz (i.e., 1 |is pulse width), the critical power gradient is 4.7x10" W/s, assuming that an optical

pulse with a wavelength of 1.55 |im is incident to a standard single-mode optical fiber with an effective cross section

area of 50 |Im^ an optical fiber length of 100 km, and the nonhnear refractive index of 3.2x10^ m^AV [7]. The

calculated critical power gradient roughly agrees with the experimental results for the 1 |is pulse width shown in Fig.

3(2). Equation (2) shows that the critical condition for the effect of SPM is more severe for a narrow bandwidth

(i.e., a long pulse width). Therefore, the critical conditions for pulse widths of 10 |is (i.e., 100 kHz receiver bandwidth) were also measured. The measured critical power gradient k was 2.1x10^ W/s and the pulse power of the

critical incident pulse amphfied with EDFAs was 16.7 dBm.

Figure 6 shows the calculated critical power for stimulated Brillouin scattering (SBS), that for stimulated

Raman scattering (SRS) using Eq. (12) in Ref.[10] and Eq. (14) in Ref. [1 1], assuming the test fiber to be a standard single-mode optical fiber with a Brillouin gain coefficient of 4.5x10"" mAV and a Raman gain coefficient of

6.5x10'"* mAV, as a function of the receiver bandwidth. The experimental critical pulse powers for SPM and FWM

are also plotted in Fig. 6. For a 1 |is or longer pulse width (i.e., a 1 MHz or narrower receiver bandwidth), the

critical pulse power for SPM was lower than those for SBS or SRS. For a 100 ns pulse width (i.e., a 10 MHz

bandwidth), the critical pulse power for FWM was also lower than those for SBS or SRS.

We also demonstrated the high performance of C-OTDR enhanced with EDFAs after 2'* integrations for 10

^is, 1 |is and 100 ns pulse widths, limited by the effect of SPM and FWM.

Figure 7 shows the C-OTDR traces for a 1 |is pulse width. The SWDR was found to be 39 dB. The incident

pulse had a pulse power of 20.7 dBm and a power gradient of 2.9x10" W/s, which was the critical condition for the

41
effect of SPM. The minimum power detectable by the receiver was about -85 dBm for 1 MHz bandwidth.

Figure 8 shows the C-OTDR trace for a 100 ns pulse width. The incident pulse power was 24 dBm, which was limited by the effect of FWM. The minimum power detectable by the receiver for 10 MHz bandwidth was about

-72 dBm. The SWDR was found to be about 29 dB. Figure 9 shows the C-OTDR trace for a 10 |is pulse width. The critical power gradient of the incident pulse was 2.1x10^ W/s, at which the incident pulse power was 16.7 dBm in this experiment. The SWDR was found to be

48 cB and the degradation by SPM was negligible in this case.

4. Conclusion

We have shown that the critical conditions at which the performance of C-OTDR combined with EDFAs is degraded by the effects of optical nonlinear phenomena in a single-mode optical fiber are given by the effect of SPM

and FWM. For 1 |is or longer pulse widths, the incident pulse powers were limited by the optical ft^equency shift of

the backscattered signal by SPM due to the optical power gradient of the incident optical pulse. For a 100 ns pulse

width, it was limited by FWM, which caused the incident optical pulse to interact with ASE and caused the incident

optical pulse to be converted to Stokes and anti-Stokes light. These critical powers were lower than those of SBS

ard SRS.

We also demonstrated the high performance of C-OTDR enhanced with EDFAs with 48 dB, 39 dB, and 29 dB

SWDRs after 2'* integrations for pulse widths of 10 |is, 1 |is and 100 ns, respectively, limited by the effect of SPM

or FWM.

Acknowledgements

We thank Y. Wakui and K. Ishihara for their continuous encouragement.

References.

[I] M. Tateda and T. Horiguchi, "Advances in Optical Time-Domain Reflectometry," J. Lightwave Technol., Vol. 7, No. 8, pp. 1217-1224, 1989.

[2] P. Healey, "Review of Long Wavelength Single-Mode Optical Fiber Reflectometry Techniques," J. Lightwave

Technol., Vol. LT3, No. 4, pp. 876-886, 1985.

[3] Y. Koyamada and H. Nakamoto, "High performance single mode OTDR using coherent detection and fiber

amplifiers," Electron. Lett., Vol. 26, No. 9, pp. 573-574, 1990.

[4] S. Furukawa, H. Izumita, I. Sankawa, and Y. Koyamada, "High dynamic range, low fading noise coherent OTDR

using erbium fiber ampHfier and LD temperature changing techniques," Proc. ECOC'91/IOOC'91, Mo.Cl-3, pp. 81-84, 1991.

[5] H. Izumita, Y. Koyamada, I. Sankawa, and S. Furukawa, "Coherent optical time-domain reflectometer with a 40-dB dynamic range enhanced with erbium-doped fiber amplifiers," Proc. OFC'92, WK6. pp.147, 1992.

[6] H. Izumita, Y. Koyamada, S. Furukawa, and I. Sankawa, "The performance limit of coherent OTDR enhanced

with optical fiber amplifiers due to optical nonlinear phenomena," to be published in J. Lightwave Technol., 1994.

[7] R. H. Stolen, "Nonlinearity in Fiber Transmission," Proc. IEEE, Vol. 68, No. 10, pp.1232-1236, 1980.

[8] G. P. Agrawal, Nonlinear Fiber Optics., Academic Press, 1989.

[9] R. H. Stolen and Lin Chinlon, "Self-phase-modulation in silica optical fibers," Phys. Rev. A, Vol. 17, No. 4,

pp. 1448-1453, 1978.

[10] R. G. Smith, "Optical Power Handling Capacity of Low Loss Optical Fiber as Determined by Stimulated

Raman and Brillouin Scattering, " Appl. Opt. Vol. 11, No. 11, pp. 2489-2499, 1972.

[II] Y. Koyamada, H. Nakamoto, and N. Ohta, "High Performance Coherent OTDR Enhanced with Erbium Doped

Fiber Amplifiers," J .Opt. Commun., Vol. 13, No. 4, pp.127-133, 1992.

42
Test Fiber

Computer!

Signal Pulse

Backscattered Signal

p-i-n-FET

Digital Processor {M]--[LPFh-^--p[S^

I ' ' • I

' '

I II ' '

I

' ' Pulse width: 1 jis (a) 21 dBm

Mixer

Fig. 1 Experimental set-up

0

LO: Local Oscillator

FC: 3 dB-Fused Fiber directional Coupler

LPF: Low Frequency Pass Filter A-D: Analog - Digital Converter

40 60 80 100

Fiber length (km)

120

Fig. 2 C-OTDR traces with l|J.s pulse width

for incident pulse power of (a) 21 dBm and

(b) 24 dBm.

I

' ' ' I • "

. 2dB Pulse width: 100 ns

(a) 24 dBm

40 60 80 100

Fiber length (km)

12C

Fig. 4 C-OTDR traces with 100 ns pulse

width for incident pulse power of (a) 24 dBm

and (b) 25 dBm.

116 120 124

Frequency (MHz)

116 120

Frequency (MHz)

124

(a) 21 dBm |

1.547 1.551 1.555 Wavelength (fim)

1.547

(1) incident optical pulse waveforms

1.551 1.555 Wavelength (|j,m)

(3) optical spectra

in the wavelength domain

(2) optical spectra

in the frequency domain

Fig. 3 Experimental results for 1 |is pulse width at (a) 21 dBm and (b) 24 dBm.

(1) incident optical pulse waveforms,

(2) optical spectra in the frequency domain of the transmitted optical pulse,

(3) optical spectra in the wavelength domain of the transmitted optical pulse.

43
(a) 24 dBm

-100 ns

|ToOmW

100

(b) 25 dBm

120

Frequency (MHz)

1.547 1.551 1.555 Wavelength (|j.m)

100 mW

—100 ns-

(1) incident optical pulse waveforms

120

Frequency (MHz)

(2) optical spectra

in the frequency domain

140 1.547 1.551 1.555 Wavelength (|i.m) (3) optical spectra

in the wavelength domain

Fig. 5 Experimental results for 100 ns pulse width at (a) 24 dBm and (b) 25 dBm.

(1) incident optical pulse waveforms,

(2) optical spectra in the frequency domain of the transmitted optical pulse,

(3) optical spectra in the wavelength domain of the transmitted optical pulse. 35

pulse power for SRS^

I 20

a 15 5

I I I ^ y I I rr Calculated criticalI III 1

Calculated critical pulse power for SBS

• Measured critical pulse power for SPM

Measured critical pulse power for FWM

' I I t 1 1

1 |is pulse width j 2 ^ ^ integrations

lOL0.1 1 10 Bandwidth (MHz)

Fig. 6 Critical pulse power of the nonlinear

phenomena as a function of bandwidth.

0

20C

-lot

' I

'

'

' I' ' ' ' 100 ns pulse width

integrations

100 150

Fiber length (km)

Fig. 7 C-OTDR trace for 1 |.is pulse width

at 21 dBm.

0 ' ' ' '

I

' > > • I 10 \is pulse width

2^^ integrations

50 100 150 200

Fiber length (km)

Fig. 8 C-OTDR trace for 100 ns pulse width

at 24 dBm.

100 150 200

Fiber length (km)

Fig. 9 C-OTDR trace for 10 ^is pulse width.

44
WINDOWING TECHNIQUE FOR ACCURATE

MEASUREMENT OF LOW REFLECTANCES BY OTDR

L. DUCOS and P. FACQ

LCIE - 33, avenue du g^n^ral Ledere - 92260 FONTENAY AUX ROSES - FRANCE

** IRCOM - 123, avenue Albert Thomas - 87060 LIMOGES Cedex - FRANCE

INTRODUCTION

Accurate measurement of low reflectances by OTDR is essentially a matter of

1) correct discrimination between two contributions overlapping over the whole deadzone length on

the OTDR display: the power R(t) returned by the reflective element under measurement and the power S(t)

backscattered by the fiber portion corresponding to the deadzone length.

2) appropriate signal analysis in order to reduce the final uncertainty on the measured reflectance.

On the basis of the discrimination-integration method reported by Blanchard et al [I], we propose

an improved signal processing technique leading to a better accuracy on measured reflectance values.

REFLECTANCE MEASUREMENT BY OTDR

The reflectance R(dB) corresponding to a reflective element can be written as [1]

R(dB) = 10 log,o{K EJP,]

where E^ is the reflected pulse energy, and is the backscattered power value at the reflective element

abscissa. The backscatter parameter K, with the dimensions of frequency, can be reached with the help of

reflectometric measurements [2]. Blanchard et al have already shown in reference [1] that when an OTDR

is used to measure the reflectance of a refractive index discontinuity (connector, coupler, splice,..) on an

optical fiber link, the reflection coefficient can be written as

R = f''[B(t) - S(t)] dt = K |£ (1)

where B(t) and S(t) are, respectively, the total backward signal and the backscattered

contribution, to is time corresponding to the

reflective element abscissa and tj is time

depending upon the OTDR dead zone. We set T= tj-tQ . The determination of the signal S(t)

has been shown in reference [1].

Figure 1 : evaluation of the reflected energy

45
DEFINITION OF THE REFLECTION COEFFICIENT

The reflection coefficient expressed by relation (1) can be expressed as the ratio of the reflected pulse

energy to the incident pulse energy Ej in relation with the considered reflective event:

f'[B{t) -S(t)] dt R,{t) dt

S(to)/K ft^ ^ r ^^^^^ dt

where R^{i) and Rj(t) are, respectively, the reflected and the incident signals in relation with the reflecting event. Rj(t) can be determined with knowing the pulse shape of the used OTDR and the energy S(to)/K. It

is possible to express relation (2) with the help of convolution products, considering that a limited integral on [to,ti] acts on Rfi) and Ri(t) as a low pass filter H with a rectangular impulse response h(t).

C^'"^' '^'^ R,(e) * h(e)

" /''R,(t) dt" «i(«> ^'^(S'

The uncertainty on R depends on the amount of noise n(t) added to the signals R^Ct) and Rj(t). This

uncertainty is directly related to the signal-over-noise ratio at the output of the filter H. Denoting x(t) the

input signal and assuming an additive centered noise n(t) with power spectral density No/2, one can show

that the SNR at the output of filter H is:

SNR = 4

No

where = J^^x^lt) dt is the input signal energy and

F ^ [j!:x(e) d0f (5) = T EX

is a positive factor depending on the shape of the signal x(t). With the help of Schwarz inequality, it can be

shown that 0<F< 1. F reaches its maximum value F = 1 when x(t) is rectangular.

THE WINDOWING TECHNIQUE

The OTDR impulse response (here denoted x(t)) generally departs from the rectangular shape on

[0,T]. The leading and trailing edges of x(t) take more and more relative importance with high resolution

OTDRs. So, F can fall to low values.

The noise spectral density No/2 being fixed in relation (4), one can see that SNR is directly

proportional to the product FE^.

46
A numerical study of the usual signal x(t) shows that starting to restrict this signal by right-side

windowing into a smaller interval [0,T'] C [0,T] leads to an increase of F not compensated by the

correlative decrease of E^. A resulting increase of FE^^ is observed as the restriction begins. FE^ grows as long as d(FEx)/dT' remains negative. Consequently, the windowing process is stopped when the product

F'E' of the truncated signal ^t) = x(t) rect[o,T'] (t) satisfies d(F'T')/dT' = 0. That is

Figure 2: right-side windowing Figure 3: right and left sides windowing

The right-side windowing (Figure 2) is temporarily stopped when the sample f(T') reaches half the

mean-value of ^t) on the interval [0,T']. Then, we compare the sample f(0) with half the mean-value. If ^0) > < > /2 there is no hope of amelioration of the SNR by making a left-side windowing. On the other

way, if ^0)< < >/2, we must restrict ^t) to [T",T'] (Figure 3) until

2 (T'' - H") L-'t dt <C2>

(8)

The new mean-value of f(t) is then increased on its definition interval [T",T'] in comparison with the

precedent value on [0,T']: <^2^ ^ < >. Then it is necessary to reconsider the stopping criteria for the

right-side windowing in relation with <f2> the new mean-value of f(t) and, eventually, to continue to decrease T'. This iterative process is pursued until

r(T") = f(T') = <f>/2 (9)

where <f> is the mean-value of ^t) on the domain [T",T']. Criteria of relation (9) is called optimal

windowing criteria. The corresponding amelioration on SNR, called SNA, can be expressed by:

SNA = [h'-t")

[/:

x(t) dt

x(t) dt

12 (lo:

47
APPLICATION OF THE OPTIMAL WINDOWING CRITERIA

The optimal windowing criteria has been applied on theoritical and experimental signals to determine

the SNA. Indicative values of SNA obtained for a triangular signal, a 20 ns pulse duration (OTDR A) and

a 100 ns pulse duration (OTDR B) are 0.73 dB, 1.61 dB and 2.50 dB respectively. Then, the optimal

windowing criteria has been applied in accordance to relation (3) to determine the reflectance R(dB)

corresponding to 10 log,o(R). Different selected values of R(dB) from -60 dB to -71 dB have been tested and were achieved with the help of a reflectance simulator [3]. Table 1 shows results of these measurements and

associated repeatability uncertainty for the Blanchard technique (relation(2)) and the windowing technique

applied to the same measured signals, ones referring to OTDR A and the others to OTDR B.

Blanchard

technique

(dB)

- 60.15

+ 0.54

- (il.lX

+ 0.46

- 63.91

± 0.64

- 66.32

± 0.60

- 67.65

± 0.81

- 69.23

+ 0.85

- 71.06

± 1.32

- 59.66

± 0.13

- 62.19

± 0.16

- 64.09

± 0.15

- 66.02

± 0.21

- 68.10

± 0.27

- 70.38

± 0.45

- 72.80

± 0.64

optimal

windowing

technique

(dB)

- 60.28

± 0.51

- 62.35

± 0.44

- 64.12

+ 0.58

- 66.61

± 0.55

- 68.12

+ 0.73

- 69.87

+ 0.84

- 71.86

± 1.17

- 59.51

± O.IO

- 61.99

± 0.10

- 63.90

± 0.09

- 65.82

± 0.12

- 67.90

± 0.17

- 70.27

± 0.27

- 73.04

± 0.57

Table 1: reflectance and associated repeatability uncertainty (± la) on 20 measurements for

OTDR A ( pulse duration = 20 ns) and OTDR B (pulse duration = 100 ns) at X = 1310 nm

DISCUSSION AND CONCLUSION

With the application of the optimal windowing technique in reflectance measurements, the uncertainty

due to noise and random errors is reduced by an amount which corresponds to the correlative increase of

the signal processing SNR. In the authors' knowledge, the accuracy of these measurements is the best yet

reported in the low reflectance range.

A comparison made with the matched filtering technique (optimal in the presence of additive

stationary noise) has shown that, due to the actual nature of noise, the accuracy of the experimental results

is not better.

The optimal windowing is an efficient simple-to-implement technique. It can be quickly and easily

set to work into an OTDR software.

REFERENCES

[1] P. Blanchard, P.H.Zongo, P.Facq, J.Rochereau.

Technical digest. Symposium on optical fiber measurement. Boulder 1990.

[2] P.BIanchard, P.H.Zongo, P.Facq. Electronics letters. Vol. 26, 25, pp. 2060-2061, dec. 1990.

[3] P.BIanchard, L.Ducos, P.Fournier, J.Dubard, F.Gauthier.

Technical digest. ESI Publications, OPTO 93, Paris, pp. 255-258. .

48
NOISE ERROR IN OTDR SPLICE LOSS MEASUREMENT

Francesco CAVIGLIA, Piergiorgio RICALDONE

CSELT, via G. Reiss Romoli 274, 10148 Turin (Italy)

Abstract - The noise induced error in OTDR splice loss measurements is discussed for the

classical loss evaluation procedures and for an alternative method, the last one giving occasionally

the best results.

1 - Introduction

The OTDR (backscattering) technique is very often used to measure the losses of splices (or

any other localized loss) in an optical link. The noise on the OTDR trace induces a random error on the measured losses; this error is usually reduced by increasing the averaging time and using

procedures that approximate the noisy trace by a straight line, as discussed in the following. (The

noise on the trace induces on the measured losses a systematic error as well, its magnitude too is considered below).

In this paper, the relationships between the noise observed on the trace and the error on the

measured losses are analyzed for the classical loss evaluation procedures and for an alternative

method. These relationships help in choosing the best procedure for the different conditions and

to evaluate the possible error by observing the noise on the trace.

2 - Loss evaluation procedures

The most basic method to estimate a splice loss is the two point method. The splice loss is calculated starting from the level difference between two points taken on the trace at the two

sides of the splice, corrected for the presumed fiber loss between the two points. Using this

method, the noise induced error is quite high. Therefore, the splice loss is usually evaluated by

measuring the level difference between two straight lines approximating the trace (Figure 1). Two

different methods can be used for obtaining the parameters of the approximating lines.

2.2 - Variable slope fitting

Using the classical way {variable slope method) the position and the slope of the

approximating lines are determined by the least square method, with reference to two fibre spans

(Dj and D2 ) at the two sides of the splice, selected by the instrument operator. This method is implemented on almost all the commercial OTDR

The position of point P, that is the position where the distance between the two line is evaluated, should in theory be at the center of the step identifying the splice on the OTDR trace. In the practice, point P is often located to the left of the step (at the end of D^) as its position has

usually a negligible influence on the measured loss (the error on the loss is equal to the product of

the misplacement of P for the difference between the attenuation coefficients of the two fibres).

2.3 - Fixed slope fitting

The alternative fitting method herein considered (fixed slope method) is similar to the previous

one, but the slope of the two lines (fibre attenuation coefficient) is fixed and stated by the user,

49
while just the line vertical positions are calculated by the least square method. (When Dj and D2

approach to zero, this method becomes equivalent to the two point method).

This alternative method is attractive when fibre spans and D2 are quite short and the

horizontal separation between the two line segments (d^ plus d2) is large. This can occasionally

happen in measuring splices between short fibre sections or because of the deadzone following the

splice reflection. Under this conditions the noise induced error on the estimated fibre attenuation

coefficients (line slopes) can be noticeable (see point 3.1) strongly affecting the loss measured

with the previous method. On the other hand, the fibre attenuation coefficient is very often known

with a good precision. Consequently this alternative method can offer better performances than

the classical one, as discussed further in section 3. The fixed slope method is not included within the fijnctions offered by the commercial OTDR

instruments, however it can be easily applied when useful. In fact, OTDR instruments, when using

the variable slope fitting method, indicate not just the calculated loss (Ay), but also the calculated

slopes: before the splice (a^) and after the splice (aj ). Moreover the positions of the markers

identifying the fitting zones and the point P are known. The loss (Af) given by the fixed slope

method can thus be calculated as: D] D2

Af = Av + (a 1 - aa ) ( Y+ d 1 ) + (a2 - aa ) ( Y+ d2 )

where is the fixed slope (fibre attenuation coefficient) assumed for the lines.

For: Dj=D2 = D and dj = d2 = d the above equation reduces to:

Af =Av + (-^— - aa)(D + 2d)

The above computations (or directly those for deriving the fitting line parameters) can be easily

performed if the OTDR is connected to a portable computer.

3 - Noise induced error

3 .1- Random error

Table 1 reports expressions to calculate the amount of the r.m.s loss error from the trace amplitude, for the different loss evaluation methods. Expression in Table 1 are derived assuming

that the line fitting on each segment D[ is done on N uncorrelated samples (being N= (Dj + S)/S ). In the practice, the distance S between uncorrelated samples is either the OTDR sampling

distance or the OTDR receiver rise-width, whichever is greater. (The latter can be easily observed on the trace, at the front of a non-saturating reflection).

Parameter Oa is the r.m.s. value of the possible deviation, from the assumed ,of the fibre

attenuation coefficients (as indicated by the OTDR, that is including any OTDR loss scale error).

On the graph in Figure 2, an example is given for the dependence of the error from the fitting zone length. On the same graph some data obtained experimentally are reported. The data in Figure 2 confirm the advantage of the fixed slope method when the fibre spans at the splice sides are short. In the example this advantage is evident for spans under 500 m , while for lengths over

1 km, the possible error on the assumed attenuation coefficient impairs the result.

50
Incidentally, the error (ctcJ in the fibre attenuation coefficient evaluated using the least square

method fitting on a fibre span (D) is given by (c being the r.m.s. noise on the trace):

12 D S

0^2= (c/D)2 + s) + 2 S)

for D » S: aa2^ 12(S/D)(a/D)2

while a two point evaluation trough the extreme points of the segment would give:

Ca2 = 2(a/D)2

3.2 - Systematic error The noise voltage is assumed to have a symmetrical distribution around the receiver

backscattering power. However, on the logarithmic trace displayed on the OTDR screen this

distribution is no longer symmetric and, as a consequence, the mean value on the trace does not

coincide with the trace position in absence of noise, but it is shifted toward lower levels. As the

noise (and the level shift) increases along the trace, this effect can influence the fibre and splice

loss measurements when the least square method fitting is used.

The analysis shows that the downward shift (in dB) of the mean value of the trace is nearly

equal to 0.45 o^, where a is the r.m.s. noise (in dB) on the trace. Moreover it can be shown that

the relative error in evaluating the loss (A) of a splice (or a fibre section) is again nearly

proportional to and given by: AA/A = 0.45 a^. This means that the error remains below 2% for

a below 0.2 dB (more than 1 dB peak to peak on the trace). The error always acts in increasing

the measured loss. When the noise is high, this effect can be avoided using fitting algorithms better than the simple

straight line one, e. g. that presented in [1].

4 - Conclusions

The reported equations allow to estimate the noise induced by the noise on the trace on splice

(and fibre section) losses measured by OTDR instruments, and evaluate the advantages of the

methods based on the straight line fitting. Moreover they indicate the influence of the receiving

system rise time on the noise induced error under similar r.m.s. noise condition.

The examined alternative loss evaluation method appears to be convenient in respect to the

classical one in a range of conditions that can be again deduced from the equations in Table 1.

Discussion at point 3.2 allows to estimate the systematic error induced by the noise asymmetry.

References

[1] - D. Messing, A. Luthra, "Estimating the power and decay rate of noisy optical

reflectometer return signals". Technical Digest of OFC "90, San Francisco (USA), 1990.

51
Figure 1 - Splice loss evaluation

using straight lines

fitted to the trace

Loss evaluation

method

General expression for the noise induced error Special cases

Two points

2 V 2 2^2

1=1 1-1

( P at the center of the step )

For:

2

crj= (72= O" dj= d2= d

2 2,2

2 (7 + 2 cr^ d

Variable slope

fitting

2 ( 12 (— + d. ^

itt M D.+ S D.( D. + 2S ) J

}

(P at the location actually used for loss evaluation)

For:

j

2

Dj = = D

dj.d^.S «D

D

Fixed slope

fitting

2^2S 22.Di n2

^1 » D.+ s r1 V 2 V

( P at the center of the step )

For: <

2

Dj = D2 = D

I dj ,d2,S «D

^ ^2 S 2 D * 2

(J _ r.m.s. noise (in dB) on the trace in the D zone

i = r.m.s. noise (in dB/km) between the assumed and the actual fibre attenuation coelTicients

<Tjj^ = r.m.s. error (in dB) on the measured splice loss S = distance between uncorrelated samples

It is assumed tliat (dj+d2),Dj and are always greater than S

Table 1

- Noise induced error for the different loss evaluation methods

2m

C

Two points >

(dB)

1 Figure 2 - Example of the error r.m.s. amplitude

as a function of the

fitting zone length

0.5 Ve slope

7 XTFixed slope

S = 25m

CTj = CT^ = O- [dB]

CT^= 0-/2 [dB/km]

dj= 60 m

Experimental data

(for (7 ~ 0.1 dB)

O Variable slope • Fixed slope

0.5 1.5

52
Effects of Fiber Parameter Mismatch on Uni-Directional

OTDR Splice Loss Measurement

James Warder, Ming-Jun Li, Paul Townley-Smith, and Castas Saravanos

Siecor Canadian Cable Plant

1370 Fletcher Road. Saskatoon, SK. , S7K 3L7

I. Introduction

Optical time-domain reflectometry (OTDR) is an effective tool for measuring fiber attenuation,

locating faults, and determining splice loss. It is well known that bi-directional OTDR

measurement is required to obtain accurate results [1]. This is especially true when measuring

fiber splice loss. However, in the field, it is not always practical or economical to have access to both ends of the tvv^o spliced fibers, and uni-directional OTDR measurement is used to estimate

the splice loss. In uni-directional OTDR measurement, the accuracy is affected by the directiondependent backscatter change due to the parameter mismatch of the two spliced fibers. The

OTDR signal level depends on the scattering coefficient, the group velocity which is inversely

proportional to the group refractive index, and the capture fraction which varies as a function of

the mode field diameter (MFD) and the effective refractive index of the fiber [2]. Thus, any

differences in the scattering coefficient, the refractive index profile, or the MFD of the two fibers

will introduce errors in the uni-directional OTDR measurement. To date, the fiber industry has

developed "rules of thumb" to correct errors caused by the MFD mismatch only [3-4], while the

contributions of all other components are ignored. When we splice the same fiber type, their

scattering coefficients and refractive indices are similar. As a result, the contributions of these

parameters to the backscatter change are small. But, they can be significant if different fiber

types are spliced to each other.

In this paper, v/e examine the effect of fiber parameter mismatch on the backscatter change. As is shown, changes in fiber parameters such as scattering coefficient and refractive indices, should

be considered to develop accurate rules of thumb.

II. Theory

In an OTDR, the backscattered power PjjsO) received at the launch end at time t, fi-om an input

pulse of power Pq and width Wis [2]

^*.(0 = ^o^il(z)exp(-2az) (1)

where a is the attenuation coefficient, and rj the backscatter factor. This backscatter factor is propertional to:

- the scattering coefficient a^,

- the group velocity Vg,

- the backscatter capture fraction B(z).

For weakly guiding single mode fiber, using Gaussian approximation, rj is given by [2]

11 = 0.5a,5(z)K, = , (2)

12ca,

53
where «e^is the effective refractive index, ng is the group refractive index, and c and kg are the

light velocity and wavenumber in free space, respectively. When two fibers are spliced, the splice

loss determined by uni-directional measurement is

a^^=\AS\ + ABS (3)

where A^" is the difference in OTDR signature reading (defined as SlogPjjg) at the splice point for

the two fibers, and ABS is the error introduced by the backscatter change due to the fiber

parameter mismatch:

ABS = 5(logTii-logr|2)

= 10 log

MFD,

MFD,

+ 5 log

a.a s2 + 10 log •effl neffl + 5 log

(4)

This error can be positive or negative depending on the direction of measurement. Eq. (4) shows

that the error is caused by the mismatch in MFD, scattering coefficient, effective refractive index,

and group index. These four parameters depend on the fiber type and manufacturing process. For

the same fiber type, the contributions of the scattering coefficient and the refractive index profile are very small compared to that of the MFD and can be ignored. But for different fiber types,

these two contributions are no longer negligible. They must be considered in the error correction.

III. Experimental

To examine the effects of fiber parameter mismatch on the backscatter change, we used three

types of commercially available fibers in our experiments: Dispersion unshifted fiber A (DUA)

and fiber B (DUB) and dispersion shifted fiber (DSF). We performed bi-directional OTDR

measurements on three groups of fiber splices: DUA-DUA, DUA-DUB, and DUA-DSF. From

the OTDR signatures obtained in the two directions, the backscatter change due to fiber

parameter mismatch is separated from the splice loss as:

2

where AS7 and A5'2 are the differences in OTDR signature reading at the splice point between the

two fibers measured from both directions respectively.

Figure 1 shows the measured backscatter change for the DUA-DUA and DUA-DUB splices. In

the same figure, we have also plotted a theoretical line of backscatter change due to MFD only

(solid line) and a theoretical line simulating the DUA-DUB splice using Eq. (4) (dashed line). In

the simulation, we used Marcuse's relationship between mode field radius and cutoff wavelength

[5]. We also assumed that the scattering coefficient is proportional to the Ge02 concentration in

the core, and the ratios of refractive indices are approximately equal to that of the core refractive

index. From this figure, we see that for the same fiber type , the backscatter change is mainly due

to the MFD mismatch, with the contribution of the other parameters being negligible. But for

different fiber types, the backscatter change due to the scattering coefficient and refractive index

differences becomes significant. It is worthwhile to realize that fiber with higher backscatter does

54
not imply higher propagation loss, as the later is not only determined by the scattering

coefficient, but also many other factors such as stresses, defects, and geometric imperfections in

the fiber.

For the DUA-DSF splice, the results are plotted in Fig. 2. In this case, since the Ge02

concentration is much higher in the dispersion shifted than in the dispersion unshifted fiber, the

differences in scattering coefficient and refractive indices between the two fiber types are larger,

causing the backscatter change due to these parameters to be more pronounced. The theoretical

line using Eq. (4) has a small offset of 0.07 dB from the measured points. This difference may be

due to the non-Gaussian shape of the mode field distribution and the non-step-index profile of

dispersion shifted fiber [6]. Nevertheless, the theoretical and experimental results agree well

qualitatively.

IV. Conclusion

We have studied theoretically and experimentally the effects of fiber parameter mismatch on OTDR backscatter change. It has been shown that, for different fiber types, the backscatter

change due to differences in scattering coefficient and refractive indices, as well as MFD must be

considered. This implies that, to correct the error in uni-directional OTDR splice loss

measurements, rules of thumb that take these factors into account need to be developed by the

fiber industry. Because fiber parameters, which depend on the shape of the refractive index

profile, dopant types and concentrations, vary from manufacturer to manufacturer, and are not

easily determined for each fiber type, an accurate and simple model is difficult to develop

theoretically. Therefore, an experimental approach is recommended to establish the rules of

thumb.

References:

[1] P. DiVita, and U. Rossi, "Backscattering measurements in optical fibers: Separation of power

decay from imperfection contribution". Elect. Lett., vol. 15, pp. 467-469, 1979.

[2] A. H. Hartog, and M. P. Gold, "On the theory of backscattering in single-mode optical

fibers", J. Lightwave Techno!., vol. LT-2, pp. 76-82, 1984.

[3] F. Kapron, C. Kozikowski, and T. Olson, "Novel OTDR effects in determining losses of

single-mode fibers and splices". Proceedings of35th IWCS, pp. 338-343, 1986.

[4] S. C. Mettler, "Monte-Carlo analysis of the effects of mode field diameter mismatch on

single-mode fiber splices", 8th Annual National Fiber Optic Engineers Conference,

Washington, D.C., pp. 647-663, April 21-24, 1992.

[5] D. Marcuse, "Loss analysis of single-mode fiber splices". Bell Syst. Tech. 1, vol. 56, pp. 703-

718, 1977.

[6] B. J. Ainslie, and C. R. Day, "A review of single-mode fibers with modified dispersion

characteristics", J. Lightv^ave Techno!., vol. LT-4, pp. 967-979, 1986.

55
Backscatter Variation vs. MFD Difference

-.10 0.00 .10 .20 .30

10 Log ( MFD1 / MFD2 )

Figure 1 - Backscatter variation for dispersion unshiftedfiber A spliced to dispersion unshiftedfibers A and B. For

the same fiber type (DUA-DUA), the backscatter change is due to the MFD mismatch only, while for

different fiber type (DUA-DUB), the contribution of scattering coefficient and refractive index mismatch

must be considered as well.

Figure 2 - Backscatter variation for dispersion unshiftedfiber A spliced to dispersion shiftedfiber. The backscatter change caused by scattering coefficient and refractive index mismatch is more pronounced due to the large dissimilarity ofthe twofiber types.

56
Practical Improved Splice Loss Estimation

for Single-Mode Optical Fibers

Hisashi TANJI, Seiichi ITO

Submarine Cable Engineering Department

International Digital Communications Inc.

Abstract

An accurate splice loss in single-mode fibers can be estimated in the field based on a

unidirectional OTDR measurement combined with a splicer prediction by direct core monitoring

(DCM) technique.

1. Introduction

In the estimation of splice losses in single-mode fibers by a fusion splicer with DCM system, a

large mismatch of mode field diameters (MFDs), whose effect is omitted from the esdmation [1],

might induce an erroneous prediction. On the other hand, a unidirecdonal OTDR measurement

shows a splice loss with a rather large deviation from true one due to even a small MFD

mismatch [2]. Bidirectional OTDR measurement, which may not always be done in the field,

has been needed to obtain an actual splice loss correctly. This paper offers an improved

estimation formula taking MFD effect into account in addition to a splicer indicated loss.

2. Theory

A splice loss a in single-mode fibers may be expressed as [3]

a = af+ad+o.i IdB] (1)

ay = -201og(2wiW2/ (wi2+h;22)) (2)

where ay, a^and a, are the splice losses due to MFD mismatch, core misalignment and core tilt, respecrively, and 2w\ and 2w2are the MFDs of the spliced fibers 1 and 2, respecdvely.

A fusion splicer with DCM system indicates the estimated loss [1]:

as = o.d + a, (3)

A splice loss as seen by OTDR from fiber 1 to fiber 2 is [2]

L(l-2) = a - 101og(wi/w2) (4)

The real splice loss ttR is obtained as a bidirectional average of OTDR apparent losses:

aR = (L(l-2)-hL(2-l))/2 (5)

57
Fig.l presents the true splice loss ay and OTDR apparent loss Ly caused by MFD mismatch in

the function of MFD ratio a (= 2wi/2w2). From [2], ay can be calculated from Ly without

knowing MFDs. By taking splicer indicated loss as into consideration in the calculation, the

total splice loss can be estimated as

aE = 2.1715 + L -[ 2.17152 + 4.343(L- as)r/2 (6)

where L means L(l-2) or L(2-l). Fig. 2 gives the relationships between as, L and ag in (6)

with broken lines of MFD ratio a = 1 {a^ = as), 0.9 and 1/0.9. For example, if MFDs of

spliced fibers are limited from 9 to 10 |im, the splice losses are expected to lie within the area between the two lines of a = 0.9 and 1/0.9.

3. Measurements

Splice losses at 1310 nm in the field cable joints were measured by OTDR in both directions.

Fig. 3 shows the measured and calculated splice losses in some fibers as an example. The

deviation of ag from as (y = as) in Fig. 3 implies ay, which is negligible if IL-asI < 0.1 dB or a = 0.98 to 1.02. The estimation error appears to be attributed mainly to an error in splicer

indicated loss, which could differ from actual loss with no MFD mismatch by 0.1 dB at the

maximum (average 0.00 dB, a = 0.03 dB) 1 1 1, while an OTDR reading error (usually less than

0.01 dB) may have negligible effect on the estimation due to the significant reduction in the

result. Fig. 4 demonstrates measured and estimated losses in comparison with real losses in actual 50

splices with relatively large MFD mismatches. It is obvious that the splice losses estimated by

(6) are, on the whole, more accurate than the splicer indications, in which the underesdmated

loss contribution of average 0.023 dB (Fig. 4 (c)) is considered the effect of mixed MFD

mismatches that are expected to be a = 0.9 to 0.97 (or 1.03 to 1.11) from the calculation.

4. Conclusion

The improved splice loss estimation with a useful chart of Fig. 2 is applicable to pracdcal cable

joints, where only unidirectional OTDR measurement can be conducted during cable installation.

Care should be taken that the OTDR reading as "splice loss" contains no excess loss factors such

as micro/macrobend in the vicinity of the splice. Visual inspection on the display of fusion

splicer is also important to eliminate unanalyzable losses due to any anomaly at the splice.

References

[1] N. Hakamata, et al., "Development of a Compact Fusion Splicer with Direct Core

Monitoring System SUMIOFCAS*^ TYPE 34", Sumitomo Electric Tech. Review,

No.27, pp. 37 - 41, 1988.

[2] F. Kapron, et al., "Novel OTDR Effects in Determining Losses of Single-Mode Fibers

and Splices", 35th IWCS, pp. 338 - 343, 1986.

[3] D. Marcuse, "Loss Analysis of Single-Mode Fiber Splices", B. S.T.J. , Vol.56, No.5,

pp. 703 - 718, 1977.

58
1.5

1 -

CDa

w 0.5

V)o

o 0

Q. CO -0.5 -

- 1

1

\ \

\

r ; 1 ; ! i . True loss a

f 0TDRL(1-2) -

X

V

0.8 0.9 1 1.1 MFD Ratio a = w / w

1 .2

1 2

Fig.1 Splice Loss Due to MFD Mismatch

( Assuming = a = 0 dB )

T3

o 0.2 o

TOO

O

Q. CO

0.1 -

1' ' 1

' 1 1 1

1 " i 1' '/ 1

/ -

-1—1—)—1—

1—r—1—I

\—1—1—1\—1

/a^=0.2SdB

; o.2o\i I _jo.15\( \

/ / i/ _/o.io\ \ \

O.OSX/ \ \ \

0 X- \ \ \ \

I -0.5 0 0.5 1 1.5

Unidirectional OTDR Loss L (dB)

Fig. 2 Splice Loss Estimation Chart

CDD

o

o

Q.

aQJ

TO

=3 O

CD

CO03

CD

y

0.6

0.4 -

0.2 -

0.2

0.4

L(1-2)

>^ a.

4 0Ce(2-1

1 L(2-1)

y = ftc

0.05 0.1

Splicer Indicated Loss (dB)

Fig.3 Example of Measured and Calculated Splice Losses

59
0 0.1 0.2 Splicer Indicated Lossag(dB)

(a) Correlation Between Splicer

Indicated Loss and Real Loss

o a^(1-2) A a^2-1)

Estimated Splice Loss a (dB)

(b) Correlation Between Estimated

Splice Loss and Real Loss

40

CD1 20

Z

-0.1

T r —I 1 !

n = 50

Av = -0.023 o =0.023

Min= -0.074 Max= 0.026

0.1

(c) Deviation of Splicer Indicated Loss

from Real Loss (in Fig. 4 (a))

H a^(1-2) - H a^2-1) -

40

0)E 20 -

0.1 0.1 a - a (dB)

(d) Deviation of Estimated Splice Loss

from Real Loss (in Fig. 4 (b))

Fig. 4 Comparison of Splice Losses by Splicer Indication and Estimation

with Real Loss

60
OTDR Trace Analysis

for Fibre Characteristics Evaiuation:

a "bi-directional" Approachi

M.Artiglia O.A.Chiantore ('^),G.Fem ('^).A.Rossaro (*)

C^) FOS - Fibre Ottlche Sud, Battipaglia (SA), Italy

(*) CSELT - Torino, Italy

1.0 Introduction

Optical fibres for telecommunications are nowadays produced in longer and longer lengths: this results into a

better handleability of fibre reels from cable manufacturers and a reduction in the total number of measurements

which can lead to a lack of characterization of the product. In order to avoid or reduce this drawback, a suitable method is required allowing to check the homogeneity of tlie fibre characteristics. In this paper, the approach adopted

in Europe, according to the indications of the European Telecommunication Standards Insitute (ETSI) and of the Cenelec Electronic Components Committee (CECC), is described: this allows for the measurement of attenuation

(longitudinal) uniformity; moreover, a possible definition of an additional parameter, namely the "backscattering

uniformity", is discussed. Experimental results are presented, relating the two parameters with the actual

characteristics of the fibre (e.g. attenuation and mode field diaineter).

2.0 General

The backscattering trace is known to be related to some of the most important fibre characteristics: in the hypothesys of gaussian distribution of the field into the fibre, also an analytical relationship is known [1]. However,

taking into account the backscattered power collected from just one end of the fibre introduces some uncertainties and

errors in the determination of those parameters, and it has been eventually recognized that the traces from both ends of

the fibre ("bi-directional" measurement) need to be considered; in particular it was proposed [2] to calculate the point

by point sum (S) and difference (D) of the traces, tlius obtaining, on a logarithmic scale [3]:

where

S(z)=Ki+51og(F(z)

)

D(z) =K2-az

F(z) =

^27cn(z)o)(z)^

ar(z) (2.3)

and

X wavelength [pm] n refractive index

CO mode field diameter []im]

^ aj Rayleigh coefficient [pm" ] a attenuation coefficient [dB/km]

Ki,K2 constants [dB]

Equations (2.1) and (2.2) show that it is possible to separate the contributions to the backscattered power coming

from geometrical parameters (i.e. refractive index and mode field diameter) variations (eq. 2.1) and attenuation

variations (eq. 2.2).

61
3. Attenuation Uniformity

The curve D(z), as defined in (2.2), can be approximated by a best fit straight line, the slope of which is the attenuation coefficient of the fibre. In order to check the attenuation variations along the fibre, the a value can be compared with a "local" attenuation coefficient, named a^, obtained considering just the backscattered power from a piece of fibre (window).

Allowing this window to slide along the fibre with a given step (e.g. the OTDR distance point), a new function a^f,(z)

is obtained giving an indication of the local attenuation of the fibre. The attenuation uniformity parameter can be defined as the maximum (positive) difference between avv(z)

and a. In formulae:

Xa = max{a^{z) -a) (3.1)

Xa allows to evaluate the longitudinal uniformity of tlie attenuation coefficient of a fibre; moreover it gives the "worst case" attenuation for any piece of fibre as long as the "window".

An important feature of this method is the possibility to localize and (if necessary) remove any attenuation

coefficient discontinuity [4,5].

4. Backscattering uniformity

Starting from the curve S(z), a backscattering unifonnity parameter X5, expressed in dB, can be defined as the maximum vertical distance between two points of the curve S(z). Assuming that the refractive index variations along tlie fibres are negligible then the only term depending on z

in F(z) is the MFD (co): S(z)=K3-101og(co(z) ) (4.1)

where Kj is constant with z. In order to expand this equation, we assume realistically tliat the MFD variations are small with respect to its minimum value co,y|:

a)(z)=Q)in(l+Aa)(z) /cDm) (4.2)

Aco(z)/o)m<<l (4.3)

S(z)=K4-4.3.Ao)(z)/(Om (4.4)

which in turn, assuming e.g. co,yi=10 pm, results into tJie following relation beetwen MFD variations and the backscattering uniformity parameter:

0.43*Aconiax=Xb (4.5)

This approximation was tested analysing about 100 single mode step index OVD fibres (as per ITU-T Rec. G.652) of different lengths at both 1310 nm and 1550 nm. Figure 1 shows the experimental points as well as the

straight line resulting from eq. (4.5) (theoretical line). X5 clearly overestimates the MFD variation, but, however, the

fitting error is quite low. The shift between the tlieoretical and the best fit line is probably due to the noise: a comparison between the MFD variation at 1310 nm and the X5 at 1550 nm, wliere the SNR is greater, confirms a much lower shift beetwen the two lines (Fig. 2). In order to better understand the power of X5 in detecting fluctuation of functional fibre parameters, let us

consider the Marcuse relation and the well-known cut-off wavelength expression for the step index fibres [6]:

62
co=d 0.65 + 0.434

^ 1.5

+0.0149 (4.6)

with:

C

- 2.4

-*NA (4.7)

where d is the fibre diameter, NA the numerical aperture and Xq the cut-off wavelength. Assuming d=8.4 pm and

NA=0.12 (typical values) and differentiating:

Aco=2.3*AXc (4.8)

hence:

(4.9)

where aX^ is in pm: a considerable consequence of this relation is that the X5 parameter could be adopted to verify if the propagation is really single-mode also in long fibres where the cut-off wavelength is only known (measured) at the extremities, i.e. at a distance of 25+ 50 km.

In Fig. 3 some data points are compared with eq.(4.9) (theoretical line). The same difference beetwen the two

lines and the same behaviour with respect to the SNR is observed as with the MFD variations (Fig. 4).

5.0 Conclusions

The information available from the light backscattered from a fibre has been discussed in the hypothesis of

gaussian distribution of the field: this approximation is very accurate for step index single mode (ITU-T Rec.G.652)

fibres. Starting from a "bi-directional" trace, two parameters, and X5, have been introduced; their correlation with

structural and trasmissive parameters (MFD, refractive index, cut-off wavelength, attenuation steps) has been proved

from both theoretical and experimental point of view, thus suggesting their possible use for quality control purposes on long lengths of fibres. Further study is necessary to verify if the model is applicable on fibres with non gaussian field disuibution, like SM-DS (ITU-T Rec.G.653) fibres. Finally, it is worthwile to point out that possible limitations in dynamic range and linearity of OTDRs can give

rise to incorrect interpretation as fibre defects of noise and/or equipment characteristics.

5.0 References

[1] M.Nakasawa,"Rayleigh backscattering theory for single-mode optical fibers", J.Opt.Soc.Am., Vol.73, No.9,

September 1983.

[2] P.Di Vita, U.Rossi, "The backscattering technique: its field of applicability in fibre diagnostics and attenuation measurements". Optical and Quantum El. 11, 17, 1980.

[3] CSELT Techn. Staff,

" Fibre Optic Communications Handbook". II ed., T.P.R., 1990.

[4] A.Chiantore et al., "Comparison of algorithms for evaluating fibre attenuation longitudinal uniformity", submitted

to Symposium on Optical Fiber Measurements, Boulder (CO), sep.1994.

[5] D.Cuomo, G.Ferri, "Evaluation of fibrel longitudinal unifonniiy: methods and algorithms", Proc. EFOC&N,

Heidelberg, June 1994.

[6] D.Marcuse, "Loss Analysis of Single Mode Fiber Sphces", B.S.T.J., May-June (1977).

63
64
DETERMINATION OF THE INHOMOGENEITY OF A SINGLE-MODE FIBER

ALONG ITS LENGTH

BY BIDIRECTIONAL OTDR MEASUREMENTS AT TWO WAVELENGTHS

Clemens Unger , Waldemar Stocklein

SIEMENS AG, ON NK E K 5, Austr. 101, D-96465 Neustadt, Germany

ABSTRACT

The inhomogeneity of the most important fiber parameters along the fiber length, e.g., modefield diameter (MFD), cut-off wavelength (k^), MAC-value (= MFD / X^) and zero dispersion

wavelength (Xq) was investigated by means of bidirectional OTDR measurements at two wavelengths. This technique also yields spatially resolved information on both the micro- as well as macrobending sensitivity of fibers.

INTRODUCTION

The development of new optical systems for telecommunication applications or data

transmission lead to higher requirements on the optical performance of the fiber. For instance

tolerances of the material and waveguide dispersion and the resulting zero dispersion

wavelength along the length of the fiber are important in high-capacity, long-distance soliton

transmission [1]. Unfortunately the dispersion can be measured only either as an integral value over the whole length of the fiber or on short (2 m) samples taken from the fiber ends.

Furthermore the optical attenuation is a substantial parameter. Current standard single-mode

fibers, designed for operation around 1300 nm and 1550 nm, exhibit differing bending losses

along the fiber length due to shght variations, e.g. of the core diameter [2]. The newly

introduced MAC-value, which is the quotient of the mode-field diameter and the cut-off

wavelength, is a suitable measure for the characterization of the bending sensitivity of singlemode fibers [2]. However up to now the MFD and Xq were always determined by well-known

standard techniques (Variable Aperture Method, Transmitted Power Method) on short fiber

pieces (typically 2m) from the ends of the fiber. Thus no informafion on the longimdinal

homogeneity is deducible.

Hence the nondestructive measurement of the distribution of the zero dispersion wavelength

and the MAC-value along a fiber is of special interest. It has been shown that the signatures of

optical time domain reflectometry can be used to determine the Gaussian mode-field diameter

[3] and the theoretical cut-off wavelength [4] along the length of the fiber. In this paper we present the experimental determination of the more customary Petermann II- MFD and the effective X^^^P by use of the bidirectional OTDR measurement at two

wavelengths. In combination with a recently proposed technique these results can be applied to get the longitudinal distribution of the chromatic dispersion and thus of Xq.

THEORETICAL

It is well known that the backscatter signal Pbs(z) from OTDR measurements of a fiber is proportional to the OTDR signature S(z) [3]:

P,,(z) = S(z)as{z)AzPoTsW 10 ^^'^ = -^\^^,f.MFD(zf (2).

It has been shown that the influence of the power decay io-(2 oc z/10) ^nd waveguide inhomogeneity S(z) as(z) may be separated by a bidirectional OTDR measurement. The Gaussian

mode-field diameter MFD0TDR(2) [3] and furthermore the theoretical cut-off wavelength

^^theo (z) [4] could be determined along the length z of the fiber by the following relations:

65
\ ( ^fiXJheo^x.X.] = ^r;^ ^ ===> Xt^'^'^ (Z) (3)

MFD^^eo ^a,X,,X,tneo^ A ^ 1 2j ^,r,OTDR^,,x ^^^^^^

where the assumption of an almost constant Rayleigh scattering a^Cz) and a modified relation

for the MFD (according to the Petermann II MFD, which will be discussed later) was used: (4)

0.634 + 1.619- { ^

1

1.5 + 2.879-

r X

]

6 -1.561-

r X

]

7^

J

+ 0.45 |im

We recently presented new theoretical and experimental investigations [2], which are in

contradiction to former publications, where an increasing macro- and microbending loss with

increasing MFD was noticed. The results can be correctly interpreted only by the use of both

the effective cut-off wavelength A.^^'^P and the Petermann II mode field diameter, that means

the MAC-value (= MFD / ;^c^^P)-

By adding 0.2 |im to the measured Petermann II MFD the Gaussian MFD can be

approximated. The experimental cut-off wavelength X^^^ of a matched-cladding fiber with

varying refractive index difference A and core radius a within typical manufacturing tolerances

may be theoretically calculated [2] from the theoretical cut-off wavelength A.^^^^ with an

uncertainty of 5 nm (Fig. 1). Ac^^P/nm

1300

1200

1100

A= 0.39

a = 4.05 |jm

A= 0.33 %

a = 4.25 |im

1200 1300 1400 ;^heo

/ nm

Fig. 1: Experimental cut-off wavelength

as afunction of the

theoretical cut-off wavelength

of a matched-cladding fiber

(a= 4.05 fim...4.25 [im,

A = 0.33 %...0.39 %)

The computation of the theoretical cut-off wavelength can be made by use of eq. (4) where the

result of the common formula was systematically shifted by 0.45 |im. With this offset we get

the best coincidence of the experimental and calculated MFD, where the core radius and the

relative index difference were measured by the Refracted Near-Field Technique. This

systematic shift is necessary due to a deviation of standard single-mode fibers form an ideal

step-index profile, e.g. the central dip.

The determination of the Xq^^P{z) by the method mentioned above requires the measurement

of 4 OTDR traces (every direction, X = 1300 nm and 1550 nm) and of 2 MFDs (X =1300 nm

and 1550 nm). Therefore every uncertainty of these values and of the relation between X^^^^

and Aq^^P influence the estimated experimental cut-off wavelength. Hence due to the above

mentioned uncertainties this technique is mainly apphcable to determine the longitudinal

inhomogeneity of X^^^P (uncertainty ±10 nm) and not its absolute value. Additionally the

effect of the variation of the Rayleigh scattering a^(z) has to be regarded.

A relatively new aspect of the bidirectional dual wavelength OTDR measurement is the determination of the waveguide dispersion from the wavelength dependence of the MFD [6]:

66
"V J . (5)

Therefore it should be possible to calculate the chromatic dispersion (the sum of material and

waveguide dispersion) and so the zero dispersion wavelength (Xq) with the knowledge of the

longitudinal MFD distribution along the fiber and of the slope of the dispersion at Xq:

^ A

> X Q

iz) ^0 - (6)

It is also thinkable to predict the zero dispersion wavelength Xq (z) as a funtion of X^^^P (z),

which may be determined from an OTDR measurement. Xq of an ideal step-index fiber can be

calculated as a function of X^^^P with a data scatter of ± 1 nm [6] (Fig. 2).

1300

1290

1280

Xq / nm

1 1 1

,j A = 0.33%

1 1 1 1

-a = 4.05 |im *9fS

, a = 4.25 |im

A

1 1 1 = 039

1 1 1 r*fek„ 1170 1250 /""^^^ 1330

/ nm

Fig. 2:

Zero dispersion wavelength

as a function of the

experimental cut-off wavelength

of a matched-cladding fiber

(theoretical)

(a= 4.05 pm...4.25 /am,

A = 0.33 %...0.39 %)

EXPERIMENTAL

Measurements on 33 short fiber pieces (2 m) with a Mach-Zehnder-Interferometer (MZI)

(Fig. 3) confirm the theoretical dependence of the zero dispersion wavelength and the cut-off

wavelength (Fig. 2). The systematic offset of the theoretical data to measured XQ-values of

nearly 23 nm can be explained again by deviations from an ideal step- index profile of the fiber.

1325 T

1315 -

1305

1295

Fig. 3:

Zero dispersion wavelength

as a function of the

experimental cut-ofi" wavelength

of a matched-cladding fiber

(experimental)

Xq = -0.0844 Xc 1418.6 nm (7)

1150 1250 1350 XJim

Furthermore we selected matched-cladding fibers with a directional dependence of the

attenuation value, derived from OTDR-backscatter traces, when measured from both fiber

67
ends. This typically can be related to a MFD variation along the fiber length. A different

variation of the MFD at A, = 1300 nm and 1550 nm is an indicator of a longitudinal change of

the cut-off wavelength and accordingly to Fig. 2 and 3 of the zero dispersion wavelength along

the fiber.

The chromatic dispersion and consequently the average Xq of the whole (2 km long) fiber were

determined by a Phase-Shift-Measurement (PSM, horizontal solid line in Fig. 4).

The fiber was cut into short pieces and the MFD-, X^-, MAC- and ^ig'^^^ues were measured.

Additionally the zero dispersion wavelength Xq (z) was calculated from the bidirectional dual

wavelength OTDR-traces

1) according to Eq. (5) and (6) from the estimated mode-field diameter MFD(z) and

2) according to Eq. (3) and (7) from the estimated cut-off wavelength (z).

^„ / nm

1325 T« °

1320

MFD

1315

1310

Fig. 4:

Experimental Xq -distribution

ofan inhomogeneous

matched-cladding fiber

calculated

from estimated MFD and

from estimated X^fW;

local measurements with a

Mach-Zehnder-Interferometer

(MZI) and

integral value of a PhaseShift-Measurement (PSM)

0 1000 2000

Length / m

The correspondence of the results of the local measurement (MZI) and the calculations verify

the reliability of this method. The systematic shift of Xq determined by the estimated MFD can

be explained by the uncertainty of Eq. (4), where a wavelength dependent term (instead of

0.45 nm) has to be used.

In addition the MAC-value ( = MFD(z) / A.^ (z) ) can be predicted along the length of the fiber,

which characterizes the longitudinal bending sensitivity of the fiber [2]. Therefore it is possible

to distinguish between the local attenuation increase in a fiber optic cable due to either a higher

bending susceptibility of the fiber or larger encumbrances of the cable.

REFERENCES

[1] V.A. Bogatyrev, et. al.: "A single-mode fiber with chromatic dispersion varying along the

length", J. of Lightwave Tech., LT-9, No. 5 (1991) pp. 561-566

[2] C. linger. W. Stocklein: "Characterisation of the bending sensitivity of fibers by the MACvalue", Opt. Comm., 107 (1994) pp. 361-364

[31 M.S. O'SuUivan, R.S. Love: "Interpretation of SM fiber OTDR signatures"

SPIE Vol. 661, Optical Testing and Metrology (1986) pp. 171-176

[41 K.W. Kowaliuik, J. Femer: "A technique to estimate the cutoff wavelength profile in

single-mode fibers using a switchable dual wavelength OTDR", Tech. Digest Symp. on

Optical Fiber Measurements, Boulder (1988) pp. 123-126

[5] M. Ohashi, M. Tateda: "Novel technique for measuring longitudinal chromatic dispersion

distribution in single-mode fibers". Electron. Lett. Vol. 29, No. 5 (1993) pp. 426-428

[61 CD. Hussey: "Characterization and design of single-mode optical fibres",

Opt. and Quant. Electron., 14 (1982) pp. 347-358

68
COMPARISON OF ALGORITHMS FOR EVALUATING FIBRE

ATTENUATION LONGITUDINAL UNIFORMITY

A. Chiantore, G. Ferri

FOS - Fibre Ottiche Sud, Battipaglia , SA ataly)

M. Artiglia, P. Di Vita

CSELT - Centro Studi E Laboratori Telecomunicazioni, Torino (Italy)

Abstract

The proposed algorithms to evaluate the longitudinal uniformity of attenuation in single

mode fibres have been compared by means of numerical simulations and elaboration of

experimental results taken over a sample of more than 600fibres. Reliability, effectiveness, area of application and robustness to noise of the algorithms have been investigated in

several situations ofpractical relevance, in view of their possible introduction in international

standards.

1. Introduction

Attenuation is a characteristic of optical fibre which is of paramount practical

relevance. At present only the global attenuation of a given fibre length is considered in

existing standards. Recently, since production lengths have increased and lengths of

installed fibre are decreasing as fibre proceeds towards the user, the need emerged for a

better understanding and control of the longitudinal variation of fibre attenuation, in order to avoid that pieces cut from a given fibre have attenuation much higher than average

attenuation of the whole fibre. Optical Time Domain Reflectometry (OTDR) equipment [1] seems to be a very promising tool to this purpose, since the OTDR backscatter trace is strongly related to the attenuation along the fibre. Moreover, the backscattering instruments

are widely diffused since OTDR provides a powerful, versatile and practical tool for testing

optical fibre attenuation, loss of fibre joints and for locating fibre defects or faults in an

installed fibre line. There are a number of approaches to characterise the attenuation variation along the

fibre, all based on the analysis of OTDR traces taken from both ends of the fibre [2]. These

are based on the observation, made some years ago [3], that when two such traces, one

taken from each end of the fibre, are averaged, the resulting trace is a straight line the slope

of which is the attenuation coefficient of the fibre. In view of a possible introduction of a

fibre longitudinal uniformity parameter in intemational standards, there is an urgent need to assess reliability, effectiveness and robustness to noise of the proposed algorithms. The aim

of this paper is to compare two common approaches in terms of the aforesaid requirements.

The first approach, known as 'sliding window' method (SW), is the one accepted as a test method for attenuation uniformity in the European Telecommunication Standard Institute

(ETSI) and in the Cenelec Electronic Components Commitee (CECC). The second is the

approach developed as a national standard in the Electronic Industries Association (EIA),

known as 'generalised sliding window' method (GSW) [4]. Results of a comparison of the

two methods performed on special cases of practical relevance, such as fibres with splices or

attenuation discontinuities, and on the data measured on 600 standard fibres without such

non-uniformities, are presented and discussed in order to identify areas of applicability,

advantages and drawbacks of the two techniques.

2. Definitions of Longitudinal Uniformity of Attenuation

The basic principles of the two studied approaches are now briefly discussed.

The SW approach is aimed at the evaluation of the optical uniformity of the fibre by

separating its two basic components: effects due to variation of MFD and refractive index

and effects due to attenuation variation. The former are described by means of the

parameter Xt, called backscattering uniformity, which is outside the scope of this paper;

the latter are described by the parameter Z^, called attenuation uniformity. This parameter

is obtained (together with Xb ) fi"om the analysis of the OTDR traces taken from the two

69
ends of the fibre under test, and is defined as the maximum over the whole fibre length of

the function:

UA(xi)=[ai(xi)+a2(xi)]/2 (1)

where xi is the position along the fibre axis of the i-th sample of the OTDR trace,

and ai(xi) and a2(xi) are the attenuations of a fibre segment of length La centred on Xi as seen from the two ends of the fibre under test. UA(x) is a measure of the averaged local

attenuation at x and La is the length of the sliding window over which local attenuation is evaluated. The parameter is tiierefore the maximum attenuation coefficient obtainable

when cutting a piece of fibre of length La. The sliding step and La and are chosen

according to the most suitable requirements.

On the other hand, the GSW approach is aimed at evaluating only attenuation

uniformity of the fibre by means of a parameter called maximum excess loss (MEL) and

is able to predict the maximum attenuation that may occur when a piece of fibre is cut

from the fibre length under test. In this approach the maximum attenuation coefficient of

any fibre segment of length L is defined as:

amax(L) ^ Oirei +£re^L (2)

wherein aref is the attenuation coefficient of the whole fibre length without nonuniformities and Eref is the MEL. These parameters, computed by means of a suitable

binary search algorithm from measured OTDR traces, shall be given by the manufacturer

for each fibre length [4].

3. Results and Discussion

The effectiveness of SW and GSW methods has been investigated both by means

of numerical simulations and experimentally. In the tests which will be discussed, the

local attenuation coefficients of Eq. (1) were calculated either by best fit procedures, or by difference between the two ends of the sliding window (two-points method). For the

seek of brevity, only the results obtained with the two points method will be given in the

following. The sliding window width assumed throughout is La =1 km.

The first type of test which has been considered is the evaluation of attenuation in a

fibre length with an attenuation discontinuity (i.e. a splice). In this test particular

emphasis has been given to the dependence of the various attenuation uniformity

parameters on the background noise. Typical results of numerical simulations on a 7 km

fibre piece with an attenuation jump of 0.025 dB and attenuation coefficient of 0.34

dB/km, are given in Figs.l and 2. Gaussian noise with zero mean and variable standard

deviation was added to the calculated data in order to simulate measurement noise. Figure

1 shows the behaviour of the estimated attenuation versus length as given by the two

methods assuming a SNR of 1 1.8 dB, compared with the actual attenuation of the fibre.

Dependence of MEL and Xa on noise power is shown in Fig. 2. It appears clearly that

both parameters give a quite good estimate of the attenuation discontinuity even if MEL

increases faster with noise power than Xa does. These features have been experimentally

confirmed by processing the OTDR traces measured on a very uniform fibre spliced

again after being cut. Both methods tend to overestimate the attenuation value with lower

signal to noise ratios. In a second type of test multiple attenuation discontinuities were considered to investigate the variation of the attenuation uniformity parameters with the relative distance

of the discontinuities. Numerical and experimental tests on a fibre with two attenuation

discontinuities (0.05 and 0.1 dB respectively) have been performed: i) by varying their

distance and keeping fixed the optimisation interval of GSW, ii) by varying the

optimisation interval and keeping fixed the distance between the defects. It has been

observed that the MEL behaves differently for varying distances between the defects.

When the distance is lower than the harmonic mean of the optimisation interval (in this case fibre lengths from 1 to 3.5 km were considered, leading to an harmonic mean of 2

km), the MEL gives an estimate of the sum of the discontinuities; when distance is

70
longer, the MEL gives an estimate of the largest discontinuity, as Xa does (provided La

is is shorter than the distance between defects).

Finally, in order to test the performance of the methods in practical conditions, the

SW and GSW algorithms have been compared using OTDR traces taken at 1310 nm and

1550 nm on a sample of 600 fibres, which comply to factory quality assessment criteria.

The results are summarised in Figs. 3 to 6. Figure 3 and 4 show the correlation plots of

the obtained values of MEL andX^ on the whole fibre sample at 1310 nm and 1550 nm

respectively. Figures 5 and 6 show the correlation plots of the obtained values of

maximum attenuation of a 1 km fibre piece predicted for each fibre of the sample by SW

method and GSW method at 1310 nm and 1550 nm respectively. It is apparent that the

two methods are mutually consistent

It should be noted that the SW method (contrary to GSW ) enables to localise the

defects of the fibre, and therefore could be particularly useful for quality control in the

fibre factory as well as for cable manufacturers and installers.

4. Conclusions

An extensive comparison between the algorithms for evaluating the fibre

attenuation uniformity considered in EIA (GSW) and ETSI (SW) has been carried out on

the basis of numerical simulations and elaboration of experimental data obtained on a sample of more than 600 optical fibres. The tests that have been carried out show that the

GSW method and the SW method give on average the same results for fibres without

attenuation discontinuities. When attenuation discontinuities, such as fibre joints, are present, the SW uniformity parameter Xa gives a good estimation of the maximum

attenuation jump and is able to localise the defects. On the other hand, the GSW method

yields a robust evaluation of the worst case attenuation for every fibre sub-length,

providing to the system designer a good estimate of the maximum attenuation that can be

expected when cutting from a bobbin a piece of given length. If more than one

discontinuity is present in the fibre the GSW uniformity parameter, the MEL, behaves

differentiy depending on the distance between the discontinuities; in addition, GSW

method is unable to localise possible defects into the fibre, which turns to be drawback

from the point of view of installers and manufactures of cables and fibres. On the other

hand the SW method is unable to provide maximum values for attenuation when defects

are present in the fibre.

References

[1] CCnr (now ITU-T) Recommendation G. 650

[2] R.G. Gravely III, S.R. Stokes, E.L. Buckland: "Fiber optic measurement

techniques for improved product and system quality control", Proc. of IWCS '93,

p. 833

[3] P. Di Vita, U. Rossi: "The backscattering Technique: its field of applicability in

fibre diagnostics and attenuation measurements". Optical and Quantum El., vol. 1 1, p. 17- 22(1980)

[4] T.A. Hanson: "Attenuation Uniformity Characterisation", proceedings of

OFMC93, Torino (Italy), Sept. 21-22, pp. 41-44

71
72
1 m spatial resolution measurement of distributed

Brillouin frequency shift in single-mode fibers

T. Horiguchi, T. Kurashima, and Y. Koyamada

NTT Telecommunication Field Systems R&D Center

Tokai-mura, Naka-gun, Ibaraki-ken, 319-11 Japan

Abstract

The Brillouin frequency shift can be applied to measure distributed strain and temperature. We

describe high spatial resolution measurements of the Brillouin frequency shift in single-mode

fibers. We have achieved 1 m spatial resolution by using an 1 1 km fiber.

1. Introduction

Changes in the frequency of stimulated Brillouin scattering can be used to measure changes in

strain and temperature in an optical fiber. Brillouin optical fiber time domain analysis (BOTDA)

can be used to measure local changes in strain and temperature along a single-mode fiber [1]. In BOTDA, a pump pulsed light and a counterpropagating continuous-wave (cw) light are launched

into a single-mode fiber so that they interact through stimulated Brillouin scattering, and the

increase in the cw light intensity due to the Brillouin gain is measured as a function of time with

OTDR. The first measurements of distributed strain and temperature using BOTDA were

demonstrated with a spatial resolution of 100 m [1]. Since these demonstrations, many efforts

have been made to improve on this value. Recently, a spatial resolution of 5 m has been achieved

[2] by using 50 ns pulsed light which is obtained by modulating the output of a 1 .32 jj.m

wavelength cw diode-pumped Nd:YAG ring laser with an acousto-optical modulator (AOM). The

pulsed light width of 50 ns is around the acoustic phonon lifetime in silica optical fibers. If we use

a pulsed light whose width is around or shorter than the acoustic phonon lifetime, the measured

Brillouin gain bandwidth increases from its intrinsic value. This means that a higher signal-tonoise ratio is required in order to measure accurately the center frequency of the Brillouin gain

band, i.e. the Brillouin frequency shift [1]. If, in addition, we take the interaction length of the

pulsed and cw lights into account, the increase in the Brillouin amplified cw light power, i.e. the

BOTDA signal power, decreases when the width of the pulsed light decreases. These facts make it

difficult to achieve a spatial resolution of better than 5 m. In this paper we show that 1 m spatial

resolution is possible by using 10 ns pulsed light which is amplified by Er-doped fiber amplifiers

(EDFAs).

73
2. Experimental setup

The configuration of the BOTDA system used in our experiments is shown in Fig. 1 . The system

is composed of a pump-pulsed light source, a probe cw light source and a 100 MHz bandwidth

detector followed by a digital processor that performs signal averaging and analysis. The pumppulsed light source is a combination of a 1.55 \im DFB-laser diode (LD) and an electro-optical

modulator (EOM) with two erbium-doped fiber amplifiers (EDFAs). An optical bandpass filter (BPF) and an acousto-optical modulator (AOM) reduce amplified spontaneous emission noise from

the EDFAs in the frequency and the time domain, respectively. The AOM also improves the

extinction ratio of the EOM. The pump-pulsed light has a pulse width ranging from 10 ns to 1 )is and a peak power of about 20 dBm in the test fiber which induces Brillouin gain. The dependence

of the Brillouin gain on the polarization was decreased by rotating the polarization of the pulsed

light with a Faraday cell (FC). We measured the temporal increase in the probe cw light power by

integrating 2^ signals for one frequency difference between the pulsed and cw lights. This

measurement was repeated by changing the frequency difference with an interval of 10 MHz

around the Brillouin frequency shift of about 1 1 GHz at a wavelength of 1 .55 |Lim. We determined

the Brillouin frequency shift as the center frequency of the Brillouin gain band. Then, we obtained

a map of the Brillouin frequency shift along the fiber length. Measurements with the BOTDA

system were carried out on an 11 km single-mode fiber composed of two kinds of fibers, type A

and type B, as shown in Fig. 2. 10 m, 5 m, 2 m and 1 m fiber spools were fomied from the A

type fiber, while the rest was the B type fiber. The difference between the Brillouin frequency

shifts of the two types of fiber is 50 ± 1 MHz.

3. Results and discussion

The distributions of the Brillouin frequency shift along the 1 1 km fiber were measured by using 60

ns and 10 ns pump-pulsed light. The results are shown in Fig. 3(a) and (b), respectively. A local

change occuiTcd in the Brillouin frequency shift in the 10 m fiber spool as expected and can be

clearly seen in Fig. 3(a). The measured Brillouin frequency shift agrees with that of the long A

type fiber and the measurement error is better than ±5 MHz. As the strain and temperature

coefficients of the Brillouin frequency shift are 5.0 MHz/lO-^-strain and 1.0 MHz/K at 1.55 p.m,

this frequency error con^esponds to a strain and temperature measurement error of ±10"^ and ±5 K,

respectively. Figure 3(a) also shows that a change in the Brillouin frequency shift can be detected even in 5 m, 2 m and 1 m spools, although it is smaller and the spatial resolution is 6 m. The

measured change in the Brillouin frequency shift of 1 m fiber spool is 0.1 -times the real value.

This is because the Brillouin gain spectrum of the 1 m fiber spool which is measured by using 60

74
ns pulsed light is a superimposition of the spectra corresponding to the 1 m A type fiber and the 5

m B type fiber [3,4]. If we resolve the two spectra, the real Brillouin frequency shift of the 1 m

fiber spool can be obtained, although we have not tried it. Figure 3(b) shows the results of a measurement using 10 ns pulsed light. In contrast to the

result for 60 ns pulsed light, the Brillouin frequency shift in the 1 m spool can be clearly seen and

as expected is in good agreement with that of the long A type fiber. The measurement error is

better than ±10 MHz, which corresponds to a strain and temperature measurement error of +2 x 10*^ and ±10 K, respectively. This measurement error can be decreased by integrating more

signals, although the measurement time increases.

Figure 4 shows the Brillouin gain bandwidth dependence on the pulsed light width. It is seen from Fig. 4 that the measured bandwidth for a pulsed light width of 1 0 ns is two times that for

1 |is. This is because the spectral width of the 10 ns pulsed light exceeds the intrinsic Brillouin

gain bandwidth of silica optical fibers. This increase in the Brillouin gain bandwidth reduces the

measurement accuracy because the variation of the measured Brillouin frequency shift is proportional to the Brillouin gain bandwidth [1]. In addition, the BOTDA signal power decreases

as the pulsed light width decreases. On the basis of the measured results shown in Fig. 3(b) we

believe that in spite of these difficulties 1 m spatial resolution is possible.

4. Conclusion

A high spatial resolution BOTDA system has been constructed which uses a combination of an EOM and an AOM to produce 10 ns pulsed light with a wavelength of 1.55 [ivn and EDFAs to

boost it. Experiments confirmed that 1 m spatial resolution is possible with a measurement

accuracy of better than ±10 MHz over an 1 1 km single-mode fiber.

References

[1] T. Horiguchi, T. Kurashima and Y. Koyamada, "Measurement of temperature and strain

distribution by Brillouin frequency shift in silica optical fibers," in Proc. Distributed and

Multiplexed Fiber Optic Sensors 11, SPIE vol.1797, pp.2- 13, 1992.

[2] X. Bao, D. J. Webb and D. A. Jackson, "32km Brillouin loss based distributed temperature

sensor," in Proc. 9th Optical Fiber Sensor Conference, 1993, PD-3.

[3] T. Horiguchi, T. Kurashima, M. Tateda, K. Ishihara and Y. Wakui, "Brillouin characterization

of fiber strain in bent slot-type optical-fiber cables," J. Lightwave TechnoL, vol.10, no. 9,

pp.1 196-1201, 1992.

[4] X. Bao, D. J. Webb and D. A. Jackson, "Temperature non-uniformity in distributed

temperature sensors," Electron. Lett., vol.29, no.l 1, pp. 976-978, 1993.

75
TIMING CONTROLLER

( PUMP

DFB-LD

1 .55 fjm

ISO

SINGLE-MODE FIBER

EOM Fc"[|[^>l["iPF]-[>-|AOM

/FIBER

AMPLIFIER on ISO

( PROBE

)

_ DFB-LD

1 .55 \im

FILTER

PUMP-LD

1.48 nm

ISO ISO

DETECTOR

DIGITAL DISPLAY PROCESSOR

Fig. 1 Configuration ofBOTDA system.

20 -

1000

Li. o

100

80

60

^1 20 < =d ma: 0 2 m

1 1 1 1

1 I

1

1 I I r

10m

I hi

5m 2m 1m

I 1 1 n

-2010.85 PULSED LIGHT WIDTH (ns) 10.90 DISTANCE (km)

10.95

Fig. 4 Relationship of pulsed light width

and Brillouin gain bandwidth (FWHM)

(b)

Fig. 3 Measurement of distributed

Brillouin frequency shift by using

(a) 60 ns pulsed light and (b) 10 ns pulsed light

76
THE DEVELOPMENT OF AN OPTICAL FIBRE ATTENUATION

STANDARD

N R Haigh^ and T C E Jones^

^ Communications Cables Division, Helsby Technology Centre, BICC Cables Utd, Helsby,

Warrington, Cheshire WA6 ODJ, UK

^ Division of Quantum Metrology, National Physical Laboratory, Queens Road, Teddington,

Middlesex TWll OLW, UK

1. Introduction

Measurement of the spectral attenuation of single mode optical fibre is crucial to fibre

manufacturing, fibre cabling and fibre installation processes. It is the principal mechanism

through which product quality is determined and is used to identify degradation of

performance in installed all-optical telecommunication links. Two measurement techniques

are used widely to determine the attenuation of single mode fibre. These are the cut-back

and the OTDR ( optical time domain reflectometry ) techniques. The equivalence of the

two techniques has been demonstrated [1], but the artefact or method for calibrating an OTDR is still under consideration.

Fibre manufacturers seek traceability for fibre attenuation measurements in order to demonstrate the validity of performance specifications attached to their products. Indeed,

many national network providers require demonstration of traceability before fibre can be

used in their networks. Traceability is obtained through National Reference Facilities for

cut-back measurement, which are maintained by National Standard Laboratories or other

organisations with responsibility for maintaining and disseminating attenuation

measurements for their country.

International standardisation bodies have placed considerable emphasis on normalising the

methods used to determine fibre attenuation [2,3,4]. A number of National Standards

Laboratories have set up facilities in accordance with these test methods to provide

traceable measurements. However, similar effort has not been applied to the development

of transfer standards in order to transfer the accuracy obtainable in National Standards

Laboratories to the industrial user. In general, measurements are performed on fibres

supplied on standard shipping reels. The measured attenuation of fibre under these

circumstances can be highly sensitive to extrinsic factors such as temperature cycling and

vibration. Such sensitivity can be increased in a fibre which is not well selected or that is badly wound on the shipping reel, as depicted in figure 1.

NPL in conjunction with BICC Cables Ltd have undertaken a programme of work to develop a 'golden fibre' attenuation transfer standard that can be readily packaged and

transported world wide and reliably used for the calibration of spectral attenuation benches

and optical time domain reflectometers.

77
2. Detail

A specification for the fibre transfer standard was formulated along with a design approval

test programme that fully qualifies its performance. This testing programme included a

rigorous set of environmental tests comprising of the following: Dry heat, 21 days at 70 °C; Damp heat 21 days at 40°C/93% Relative humidity; Temperature cycling, 10

cycles of - 10/70 °C, Cold storage, 10 days at -40 °C; vibration testing.

The fibre was packaged so that the fibre is protected from the rigours associated with

transit, and to avoid any losses due to macrobending. In addition, the transfer standard was

designed to facilitate ready access to both ends of the fibre in order to accommodate the

destructive 2 metre fibre cut-back to be performed during the spectral attenuation

measurement procedure.

Details of the design of the fibre attenuation standard will be presented at the conference.

The most significant factor in the design of the fibre standard is related to fibre handling,

where it is essential that the fibre is free from stress along its entire length. This is considered to be vital in ensuring that environmentally induced changes in attenuation

cannot arise from effects such as tight fibre winding or fibre cross-overs. It also results in

the fibre exhibiting a very low level of polarisation mode dispersion.

3. Results

The spectral attenuation of a 6 km standard measured by the two laboratories ( NPL and

BICC ) is shown in Figure 2. The rapid increase in attenuation above 1550 nm which is characteristic of bend loss, and commonly observed for fibres wound on shipping reels

( see figure 1 ) is not observed for the transfer standards. Figure 3 shows the difference in

the spectral attenuation values as measured by the two laboratories. In the second and

third transmission windows, the difference between the two laboratories is less than

± 0.003 dB/km. However, it should be noted that a relatively large error, of the order of

0.05 dB/km occurs in the 1380 nm hydroxyl ion absorption feature of the attenuation

spectrum, where the spectral attenuation of the fibre varies strongly with wavelength. Such

an error is not unexpected and is probably indicative of a slight offset in the wavelength

calibration of the two spectral attenuation benches. Measurements of fibre attenuation were made during the above environmental test programme. Stability of the fibre attenuation

was measured to be better than ± 0.03 dB. To date, the fibre attenuation has been

measured to be stable to less than ± 0.03 dB/km during the above environmental test programme. At the end of each environmental test, the fibre spectral attenuation was

observed to return, within the measurement error, to its initial value.

4. Conclusions

A 'golden fibre' attenuation standard has been developed with an uncertainty in the total

loss of better than ± 0.06 dB ( 2a ). The short term stability of the attenuation of the

standard has been demonstrated to be better than ± 0.03 dB.

78
Acknowledgement

The assistance of P J Whitesmith, M Harrop, T Holland and M R Shaw of BICC Cables

Ltd, in the design and construction of the fibre attenuation standard is gratefully

acknowledged.

References

1. A L Vandervort, J L Swecker "OTDR attenuation measurement and effect of wind

tension", Proceedings OFME93, Turin

2. International Standard lEC 793/1, " Optical fibres - Part 1: Generic specifications"

3. CCITT Recommendation G.650, "Definition and test methods for the relevant

parameters of single mode fibres"

4. FOTP - 78, "Spectral attenuation cutback measurement for single mode optical

fibres"

79
Figure 1 : Variation of the spectral attenuation of a 12 km SMF

wound CD a shipping spool due to bending losses

Figure 2 : Spectral attenuation of 'golden fibre' standard as measured by NPL and BICC

0.700

0.600

0.500

a

O

c § 0.400

.2 -o

«-« a

se

0.300

0.200

0.100

44.

1200

++44,

-^+44.++44+

+44444 +4-

V

+

\+ v.^+^+4444444444444444.

4444444+4-

1250 1300 1350 1400 1450

Wavelength / nm

>00 1550 1600

Figure 3 : Difference between NPL and BICC measured

attenuation values for golden fibre standard

d

•is

s

0 060 T 0 050

0 040

0 030

0 020

0 010

0 000

-0 010

-0 020

1200 1250 1300 1350 1400 1450 1500 1550 1600

Wavelength / nm

80
Microwave based Measurements of Optical Fibers

Robert M. Atkins

AT&T Beii Laboratories

Murray Hill, NJ 07974

USA

Microwaves have been utilized to measure the characteristics of a carbon coating used in the

production of hermetically coated fiber, and also to measure the cross sectional area of non-circular

fibers, such as highly elliptical fibers drawn for use in cladding pumped lasers.

In the first application, absorption of microwave power by the conductive carbon coating applied to a hermetically coated fiber is used to determine the characteristics of that coating^^'. The hermetically

coated fiber is passed though a section of WR90 (X band) waveguide as shown in Figure 1. Slots

(ca 3mm x 25mm) are cut into opposing broad faces of the waveguide to permit the fiber to pass

through. One end of the waveguide is connected to a low power X band microwave source (in one

case a Gunn oscillator operating near 10.5 GHz with an output power of 10mW) and the other end

of the waveguide is connected to a diode detector. Matching screws were located at each end of the

waveguide section and adjusted to obtain the desired impedance match. The output fi^om the detector

was fed to a circuit which produced a voltage output in proportion to the change in received power

in the presence of an hermetically coated fiber. Using coated fibers whose coating conductance was

independently measured off-line, a cab"bration plot was obtained as shown in figure 2. As can be seen, there is a good linear relationship between fiber conductance and the output from the microwave

sensor. Conductance was also found to be directly related to coating thickness, through off line electron microscopy measurements. Results of an on-line test, where hermetically coated fiber was

passed through the microwave cavity while being drawn, are shown in figure 3. Deliberate changes

made in the hermetic coating process (marked "*"), reflected in later off-line measurements of fiber

resistance, are clearly shown in the output from the microwave sensor.

In the second application, the cross sectional area of a fiber was measured via the influence of the

fiber on the resonant frequency of a microwave cavity. Figure 4 shows a simple implementation of

this technique in which a resonant cavity is formed by two iris plates in a section of WR90 (X band)

waveguide. The resonant frequency of the cavity is determined by its physical length (1/2 wavelength

at resonance) and by the dielectric constant of its contents. The fiber passes through the center of the

cavity via 2 small holes (ca 3mm diameter) in the broad faces of the waveguide. The presence of a

silica fiber in the cavity raises the local dielectric constant significantly and results in a drop in the

cavity's resonant frequency. The resonant fi-equency of the cavity may easily be determined through

the use of a network analyzer. Using both single optical fibers of known cross sectional areas (determined by microscopy) and multiple fibers whose total cross sectional area was simply the sum

of the areas of the individual fibers, a calibration plot of fi:equency shift vs. cross sectional area was

obtained as shown in figure 5. As can be seen, there is an excellent correlation between the two

plotted parameters. In contrast to optical fiber diameter sensors, whose output when measuring

81
elliptical fibers is strongly influenced by rotation of the fibers major/minor axes, the output from the

microwave sensor is insensitive to the fiber orientation.

Though these two techniques are superficially similar they are based on different principles. The

former technique, for measurement of hermetic coating thickness, depends on the attenuation of

microwave power via dissipation due to dielectric heating and is, to first order, insensitive to the

microwave fi^equency. In the absence of a conductive hermetic coat, the bare fiber has little or no

effect on the transmission of microwave power through the waveguide. In contrast, the latter technique, for cross sectional area measurement, depends on the changes in resonant frequency of

the cavity due to the volume of silica dielectric (bare fiber) in the cavity and is insensitive to changes

in microwave power.

Reference

[1] US patent 5013130

OPTICAL

FIBER

MICROWAVE

POWER OUT

MICROWAVE

POWER IN

Figure 1 - Schematic of waveguide sensor for hermetic coating conductivity measurement

82
0.2 -

i 0.0 L ' i • :

.

' 0 2 4 6 8 10 12

Fiber Conductance (MHO/cm x 10 5)

Figure 2 - Calibration of hermetic coating conductivity sensor

3 I

\

'

^ 1

' 1'

\

'

1

' 1' r

01 23456789 10

Time (minutes)

Figure 3 - On-line test of hermetic coating conductivity sensor. "*" marks the change of coating

conditions. Coating resitances (measured off-line) were as follows: A - 25 kQ/cm, B- 54kQ/cm,

C - 18 kQ/cm, D - 13kQ/cm, E - lOkQ/cm and F - 24kfi/cm

83
OPTICAL

RBER

IRIS

PLATE

MICROWAVE

POWER IN

IRIS

PLATE

MICROWAVE

POWER OUT

Figure 4 - Schematic of waveguide cavity sensor for fiber cross sectional area

84
Characterization of In-fiber Bragg Gratings

K. O. HiU, F. Bilodeau, B. Male,

S. Th6riault, D. C. Johnson, J. Albert,

and T. Kitagawa*

Communications Research Centre 3701 Carling

P.O. Box 11490, Station "H"

Ottawa, Ontario, Canada

K2H 8S2

SUMMARY

Irradiation of the core of an optical fiber with ultraviolet light photoinduces a permanent

change in the core refractive index. This photorefractive effect'^^-' called "photosensitivity",

has great practical significance: it makes possible a new family of passive fiber devices that

have potential application in optical communications and optical sensors^^^. Photosensitivity

technology is reviewed in this invited paper with emphasis on the measurement techniques

used in the characterization of in-fiber Bragg gratings. The use of photoinduced in-fiber

Bragg gratings and long-pitch mode converter gratings is proposed as the basis for

measurement techniques useful for determining some of the important characteristic

parameters of optical fiber.

Characterization ofFiber Bragg Gratings

A Bragg grating photoimprinted in the core of a monomode optical fiber acts as a narrow

bandwidth reflector. The key parameters describing in-fiber Bragg reflectors are the resonant

wavelength,

'^gragg' grating length L, and the reflectivity as a function of wavelength.

From these quantities the photoinduced refractive index change 6n, and the grating coupling

constant k are derived. The grating length is normally not measured but is determined in the

fabrication process by a mask that limits the length of the fiber that is exposed to ultraviolet

radiation. Thus the principal measurement problem is a determination of the reflectivity of

the Bragg reflector as a function of wavelength.

* On leave from NTT Opto-Electronics Laboratories, Japan

85
The reflectivity of a Bragg grating is measured using standard spectroscopic techniques.

Light from a broadband light source (superradiant fiber, white light source, or light emitting

diode) is coupled into the fiber, passed through the grating and its spectral content measured

by means of an optical spectrum analyzer. The experiment gives a signal proportional to the

Bragg grating transmission as a function of wavelength of the incident light. To obtain the

grating reflection response, the experimental data is processed (normalized) to eliminate the

fiber input and output coupling losses. Because the grating spectral responses are narrowband, the signal level in the wings of the Bragg resonant wavelength can be used as the baseline in the normalization process. The equipment used in carrying out the

measurements should have a large dynamic range, ( ~ 60 dB is achievable in practice

determined by the intrinsic dynamic range of the instrumentation and the experimental setup

), in order to measure a wide range of reflectivities. The measurement of high reflectivities

(transmissions less than -60 dB) or very low reflectivities (< 4%) present special

measurement problems.

Another approach for measuring the Bragg grating reflection response is to use a single

frequency tunable laser diode. The transmission of the Bragg grating is monitored as the

diode is tuned through the Bragg resonant wavelength. Alternatively, the diode frequency can be held fixed and the Bragg resonant frequency tuned by stretching (stress tuning)

through the single wavelength of the laser. The resonant frequency of a Bragg reflector can be shifted by ~ 12 nm through strain (1%) tuning. This method is useful for measuring Bragg

reflectors having very narrow linewidths; linewidths less than 200 MHz have been measured

using this approach^^-'.

Measurement of Photoinduced Refractive Index

The photoinduced refractive index has been measured directly by irradiating one arm of an unbalanced Mach-Zehnder fiber interferometer with UV light, thereby changing its optical

path length From the results, the dispersion in the photoinduced index can be determined.

Characterization of Chirped Bragg Fiber Gratings

Photosensitivity has been used to fabricate chirped (aperiodic) Bragg gratings that have

application in dispersion compensation The gratings are chirped in the sense that the

optical pitch (i.e. effective refractive index n^^^ times the grating pitch A varies linearly along

the length of the structure. The additional key parameters describing chirped gatings are the

percentage chirp and the dispersion-delay characteristics of the Bragg reflector. The Bragg

gratings are normally fabricated to have a specified percentage chirp. However, it is useful to

relate the chirp to other characteristics of the Bragg reflector. For "weak" bandstop

reflectivity (R less than 50%) the full-width-at-quarter-maximum of the spectral response curve of a linearly chirped Bragg grating is given by the percent chirp multiplied by the

86
center Bragg resonant wavelength of the device, Xq. Alternately, an approximate theoretical

50 Xq\k\\

expression for the percent chirp is: . n^ffln[(l-R)-^]

When used as a dispersion compensator, the chirped grating is operated in the reflection

mode. Light having a wavelength that is resonant for Bragg reflection near the rear of

chirped grating is delayed with respect to light having a wavelength that is resonant for Bragg

reflection at the front of the grating. The delay dispersion characteristic of the Bragg grating

gives the relative delay for light as a function of its frequency offset from the center Bragg

resonant frequency of the chirped grating. The dispersion delay characteristic is measured

using apparatus described by K. Takiquchi et al'-^-'.

Measurement ofEffective Refractive Index of an Optical Fiber Using Photoimprinted Bragg

Gratings

Accurate values of the effective group refractive index of monomode fibers are important for

the calibration of the scale in optical time domain reflectometers (OTDRs).

The standard approach for measuring the effective group index of an optical fiber is the

pulse time-of-flight technique. The transit time for a short optical pulse to propagate over a precisely measured length of optical fiber is measured'-''-'. The technique measures the

effective group index to ± 0.0001.

In this paper we propose to measure the effective refractive index of a monomode fiber by

photoimprinting a grating in the core of the fiber using the phase mask techniquePf. The

effective index is determined by the Bragg resonant wavelength ^igragg ^'^^ phase mask

^Bragg

grating period A which are related as follows for second order reflections: ng^ = A

Since both ^g^agg ^'^^ ^ determined with a high degree of precision, the technique can

yields accurate values for n^^. However, for the measured n^^ to correspond to the effective

index of the test optical fiber it is necessary that the photoinduced refractive index be less

than 10 . From a series of such measurements of effective index at various wavelengths, the

dispersion in effective group index can be determined.

Measurement of Optical Fiber Mode Differences Using Mode Converter Gratings

The function of a mode converter is to convert light propagating in one mode of a waveguide

to another mode of the waveguide. Mode converters are fabricated by inducing a periodic

refractive index perturbation along the fiber length with a pitch that bridges the momentum

mismatch (propagation constant diff^erence) between the modes and thus allows phasematched coupling between the selected fiber modes. Photosensitivity has been used to write

spatial'-^-' and polarization mode converter gratings'^^^ in optical fibers. Mode converter

87
gratings provide a means for measuring the difference in propagation constants between two

modes in optical fiber.

To fabricate an efficient mode converter using the point-by-point writing technique it is necessary to reproduce by exposure from the side the same holographic pattern that the two

modes would trace in the waveguide core at the resonant wavelength for energy exchange

between the modes.

In the case of LPq^<-^LP^^ fiber mode converters the interference pattern between the modes

is highly blazed (because the LP^^ mode is not circularly symmetric) at an angle of a few

degrees with respect to the fiber core. In contrast, for LPqjOLPq2 mode converters the

required blaze is zero because the interfering modes are both circularly symmetric'-^^-'.

A feature of the LPqj<->LPq2 mode converter is that it has a single resonance because of the

circular symmetry of the modes and of the mode converter. In contrast the LPq^oLP^^ mode

converter has four resonances due to the 4-fold degeneracy of the LP^^ mode.

The fraction of the launch power in Mode 1 coupled by a uniform mode converter to co- propagating Mode 2 is given by

[ \K\^ + (—ff z

. 2 sm

IkI -I- ( )

where k is the coupling coefficient of the mode converter grating, z is the position coordinate

27im

measured along the length of the grating and Ap = [p, p - Pt p - ( )] with m an

integer, P is the mode propagation constant and A is the periodicity of the grating. LP

designates a linearly polarized mode of the waveguide. When phase matching occurs (Ap = 0), the total power transfer is modulated by the sin function and can reach 100 %.

Thus for a mode converter grating operating in first order, the resonant wavelength for

efficient coupling between two modes corresponds to the wavelength for which the difference

in propagation constant between the two modes is simply 27i/A.

88
REFERENCES

1. K. O. Hill, Y. Fujii, D. C. Johnson and B. S. Kawasaki, "Photosensitivity in optical fiber

waveguides: Application to reflection filter fabrication", Applied Physics Letters, Vol. 32,

No. 10, pp. 647-649, May 15, 1978.

2. K. O. Hill, B. Malo, F. Bilodeau, and D. C. Johnson, "Photosensitivity in optical fibers",

Annual Review of Material Science, Vol. 23, pp. 125-157, 1993.

3. B. Malo, K. A. Vineberg, F. Bilodeau, J. Albert, D. C. Johnson, and K. O. Hill,

"Ultraviolet Light Photosensitivity in Ge-Doped Silica Fibers: Wavelength Dependence

of the Light-Induced Index Change", Optics Letters, Vol. 15, No. 17, pp. 953-955,

September 1, 1990.

4. K. O. HiU, F. Bilodeau, B. Malo, T. Kitagawa, S. Th6riault, D. C. Johnson, J. Albert and

K. Takiguchi, "Aperiodic in-fiber Bragg gratings for optical fiber dispersion

compensation". Optical Fiber Communication Conference, OFC'94, San Jose, California,

February 20-25, 1994, Technical Digest, Postdeadline Papers, PD17, pp. 17-20, February

24, 1994.

5. K. O. Hill, "Aperiodic Distributed-Parameter Waveguides for Integrated Optics", Applied

Optics, Vol. 13, pp. 1853-1856, August 1974.

6. K. Takiguchi, K. Okamoto, S. Suzuki, and Y. Ohmori, "Planar lightwave circuit optical

dispersion equalizer", 19th European Conference on Optical Communication, ECOC'93,

Montreux, Switzerland, September 12-16, 1993, Proceedings, Vol. 3, Post deadline

papers, ThC12.9, pp. 33-36.

7. J. J. Carr, S. L. Saikkonen, and D. H. Williams, "Refractive index measurements on

single-mode fiber as functions of product parameters, tensile stress, and temperature".

Fiber and Intergrated Optics, Vol. 9, pp. 393-396, 1991.

7. K. O. Hill, B. Malo, F. Bilodeau, D. C. Johnson, and J. Albert, "Bragg gratings fabricated

in monomode photosensitive optical fiber by UV exposure through a phase mask",

Applied Physics Letters, Vol. 116: 62, No. 10, pp. 1035-1037, March 8, 1993

8. K. O. Hill, B. Malo, K. A. Vineberg, F. Bilodeau, D. C. Johnson, and I. Skinner, "Efficient

Mode Conversion in Telecommunication Fibre Using Externally Written Gratings",

Electronics Letters, Vol. 26, No. 16, pp. 1270-1272, August 2, 1990.

9. K. O. Hill, F. Bilodeau, B. Malo and D. C. Johnson, "Btrefringent Photosensitivity in

Monomode Optical Fibre: Application to the External Writing of Rocking Filters",

Electronic Letters, Vol. 27, No. 17, pp. 1548- 1550, August 15, 1991.

89
10. F. Bilodeau, K. O. Hill, B. Malo, D. C. Johnson, and I. M. Skinner, "Efficient, narrowband

LPq^<-^LPq2 mode convertors fabricated in photosensitive fibre: Spectral response".

Electronics Letters, Vol. 27, No. 8, pp. 682-684, April 11, 1991.

90
Characterization of Optical Components using

High Resolution Optical Reflectometry Techniques

W. V. Sorin, D. M. Baney and S. A. Newton

Hewlett-Packard Laboratories

Introduction : Optical time domain reflectometry (OTDR) has proven to be a very important

diagnostic tool for characterizing and trouble-shooting long haul fiber-optic communication

links. Characteristics such as the location of a fiber break, attenuation coefficient, splice loss

and link degradation can easily be measured with access to only one end of an installed fiber. Naturally, there is a desire to extend this type of measurement capability to address the area of

optical component analysis where resolution on the order of 100 |Lim or less is required.

Extrapolation of the techniques used in pulsed OTDR can become very difficult as pulse

resolution gets below about 10 mm. One problem is that the reflection sensitivity becomes

severely limited by additional noise due to the increased receiver bandwidth. A more

fundamental problem exists due to the effects of dispersion, since the short optical pulses will

need to propagate through fiber pigtails that are attached to the components being measured.

Frequency Domain Techniques : Due to the difficulties associated with pulsed approaches,

many other techniques have been investigated to realize high resolution optical reflectometry

measurements. Using the relationship between an impulse response and its Fourier transform,

high resolution reflectometry can be performed by using frequency domain techniques.

Frequency-stepped sinusoidal modulation of the optical intensity of a probe signal has been

used to achieve two-point spatial resolution on the order of a few millimeters. ^ This approach

is limited in its reflection sensitivity and dynamic range due to its use of direct detection. A

similar but more promising approach is to frequency step or sweep the optical carrier and use

coherent detection at the receiver.23

Low

Coherence

Source Test Arm

^DUT

R

A

~10n,ni

PC

Ref Arm

X

1

Figure 1 . Basic configuration used for optical low-coherence reflectometry.

91
This technique, known as coherent optical frequency domain reflectometry, can have excellent

sensitivity, dynamic range and spatial resolution. Two-point spatial resolution on the order of

tens of microns has been demonstrated.^ Presently, this technique is limited by the availability

of practical, narrow line-width, widely tunable optical sources.

Optical Low-Coherence Reflectometry : A more practical approach that does not require the

development of any new technology is the use of optical low-coherence reflectometry

(OLCR).4,5 AsimpHfied version of an OLCR is shown in Figure 1 . A continuous wave (cw)

short coherence length source, such as an edge-emitting LED, is coupled into a fiber-based

Michelson interferometer. A movable reference mirror in one arm is translated to provide a

variable optical time delay. Any optical reflection from the device under test (DUT), whose

time delay matches that from the moving reference mirror will coherently interfere to produce

an interference wave packet whose spatial extent is equal to the coherence length of the source. One inherent advantage of this technique is that dispersion broadening due to the component

pigtail and fiber interferometer can be completely eliminated by ensuring that an equivalent

dispersion is seen by the reference signal. In practice this is easily done by using the same type

of fiber in both arms of the interferometer. Other advantages are that the source does not

require any complex modulation and only a simple low frequency detector is required. This

technique has been demonstrated to achieve spatial resolution below 2 iim^ and reflection

sensitivity greater than -160 dB.^ Due to its practicality and increasingly wide use, this paper

will focus primarily on applications that have been demonstrated using OLCR.

Accurate Time Delay Measurement: One of the strengths of OLCR is its ability to measure

optical time delays very accurately. Although the measured impulse response from a reflective

surface can have a width on the order of ten microns, it is possible to determine its central

position with an accuracy of better than 0.02 |im^ (this corresponds to an equivalent time delay

of 0.07 femtoseconds). Typical applications include trouble-shooting and quality control — for

example, to determine the position of optical components within packaged fiber pigtailed

devices such as isolators, laser diodes, and photoreceiver packages. Other applications include

accurate characterization of time delays in interferometric optical circuits such as Fabry-Perot

cavities, etalons, and Michelson and Mach-Zehnder interferometers. ^ OLCR has also been

used to accurately measure silicon wafer thickness. ^ ^

The time delay measured using OLCR (as in any reflectometer) is the time it takes for an

optical pulse to propagate through the test medium. This delay is equal to the product of

physical length and group index and should not be confused with the optical path difference

which is the product of physical length and refractive index. In many situations, the parameter

of interest is either physical length or group index. Being able to independently measure one of

these parameters allows for the solving of the other, but in cases where an independent

measurement is difficult it may be possible to solve for both parameters simultaneously using

only OLCR measurements. ^ 1 Besides measuring isotropic material which are characterized

with a single time delay, differential time delays associated with transmission through

birefringent material can also be measured. This allows the difference in group velocities to be

measured in materials such as polarization maintaining fiber, 1^ rutile crystals, and LiNb03

92
waveguides.^ Using a modified OLCR, it is also possible to measure polarization mode

dispersion in long optical fiber links.

Backscatter Measurement: In addition to precise time delay measurements, OLCR also

provides reflectivity information. Due to its coherent detection scheme, reflection sensitivities as low as -161 dB have been demonstrated.^ With such high sensitivities, the -120 dB

Raylei gh backscatter reflectivity from a 10 |im section of standard single-mode fiber can be

measured easily.^' Using a high sensitivity OLCR, backscatter and waveguide imperfections

have been measured in glass waveguides, LiNb03 waveguides, GaAs waveguides, erbium

fiber, fused directional couplers and tight fiber bends.^'l^ It should be pointed out that the

Rayleigh backscatter signal obtained using OLCR has much more amplitude uncertainty than

that obtained using conventional pulsed OTDR methods. amplitude uncertainty,

sometimes referred to as coherence fading, occurs when the spatial resolution is comparable to the source coherence length. In conventional pulsed OTDR measurements, the source coherence length is usually much smaller than the pulse length so coherence fading is not a problem. Besides measuring backscatter from optical waveguides, OLCR applications also

include analytical and biomedical applications to determine scattering and attenuation

coefficients in liquid suspensions and biological tissue. ^^'^^

Return Loss Measurement: When making return loss measurements on optical components,

OLCR has several advantages over the conventional technique of using a cw source and

measuring the total back-reflected power with a standard power meter. The most obvious

advantage is that individual reflections can be spatially resolved. This can be very useful when

minimizing the return loss from components that contain many internal reflective interfaces.

Another advantage is the ability to measure reflectivities below -70 dB. This becomes very

difficult using the cw power meter approach since the background reflected signal due to Rayleigh backscatter in a standard single-mode fiber is about -70 dB per meter. Sensitivities

below this value are needed to characterize components such as angle polished connectors and

high return loss isolators.

Gratings. Optical Sources and Dispersion: High resolution reflectometry has also been used to characterize photoinduced reflection gratings within single-mode fibers. ^0 Information such as the spatial dependence of the reflection coefficient and the grating length can be independently

measured.20521 OLCR can also be used to probe active devices such as EELEDs, laser diodes

and DFB lasers. ^^22,23 Typically the semiconductor device has to be pumped to improve its transparency. In these cases, the amplitude of the reflected signal also depends on the optical

gain it experiences through the device. Chromatic dispersion can also be measured using

OLCR. This can be accomplished using the phase information within the received interference

wavepacket. Due to the wide optical bandwidth of the probe signal, accurate dispersion

measurements can be made using relatively short optical delays. This technique has been used

to measure how the dispersion through an erbium doped fiber amplifier changes as a function of

optical pump power.24

93
Increasing Distance Range : One drawback of OLCR is that long scan ranges are difficult to achieve. This occurs since a reference mirror must be physically translated to provide the

required optical delay. In practice, this makes scan ranges larger than about 0.5 meters

difficult to achieve. Recently, a new technique has been demonstrated that can increase the scan range by over 2 orders of magnitude. This is achieved by adding a recirculating optical

delay into one arm of the OLCR. By extending the measurement range, new applications for

OLCR will become practical.

Summary : Characterizing and trouble-shooting fiber-pigtailed components is often a very

difficult task. Due to packaging constraints, visual inspection is not usually practical.

Transmission characteristics, when obtainable, often give valuable information but in many

cases do not provide enough information to characterize an optical device. High resolution

optical reflectometry, namely OLCR, has become a very important diagnostic tool in many of

these cases.

References

1 . D. W. Dolfi, M. Nazarathy and S. A. Newton, "5 mm resolution optical frequency domain

reflectometry using a coded phase reversal modulator." Opt. Lett., vol. 13, p. 678, (1988).

2. W. V. Sorin, D. K. Donald, S. A. Newton and M. Nazarathy, "Coherent FMCW

reflectometry using a temperature tuned Nd:YAG ring laser," Photonics Tech. Lett., vol.

2, p. 902 (1990).

3. U. Glombitza and E. Brinkmeyer, "Coherent frequency-domain reflectometry for

characterization of single-mode integrated-optical waveguides," J. Lightwave Technol.,

vol. 11, p. 1377(1993).

4. R. C. Youngquist, S. Carr and D. E. N. Davies, "Optical coherence-domain reflectometry:

A new optical evaluation technique," Opt. Lett., vol. 12, p. 158 (1987).

5. B. L. Danielson and C. D. Whittenberg, "Guided-wave reflectometry with micrometer

resolution," Appl. Opt., vol. 26, p. 2836 (1987).

6. X. Clivaz, F. Marquis-Weible and R. P. Salathe, "Optical low-coherence reflectometry

with 1.9 |Lim spatial resolution," Electron. Lett., vol. 28, p. 1553 (1992).

7. K. Takada, T. Kitagawa, M. Shimizu and M. Horiguchi, "High-sensitivity low coherence

retlectometer using erbium-doped superfluorescent fibre source and erbium-doped power

amplifier," Electron. Lett., vol. 29, p. 365 (1993).

8. B. L. Danielson and C. Y. Boisrobert, "Absolute optical ranging using low coherence

interferometry," Appl. Opt., vol. 30, p. 2975, (1991).

9. Unpublished measurements performed by authors at Hewlett-Packard Laboratories.

10. C. W. Lawson and R. R. Michael, "Fiber optic low-coherence interferometry for noninvasive silicon wafer characterization," Journal of Crystal Growth, vol. 137, p. 37,

(1994).

11. W. V. Sorin and D. F. Gray, "Simultaneous thickness and group index measurement

using optical low-coherence reflectometry," Photonics Tech. Lett., vol. 4, p. 374, (1992).

94
12. M. Tsubokawa, T. Higashi and Y. Sasaki, "Measurement of mode couplings and

extinction ratios in polarization-maintaining fibers," J. Lightwave Technol., vol. 7, p. 45

(1989).

13. N. Gisin, J-P. Von der Weid and J-P. Pellaux, "Polarization mode dispersion of short and

long single-mode fibers," J. Lightwave Technol., vol. 9, p. 821 (1991).

14. W. V. Sorin and D. M. Baney, "Measurement of Rayleigh backscattering at 1.55 [im with

a 32 |j.m spatial resolution," Photonics Tech. Lett., vol. 4, p. 374, (1992).

15. K. Takada, "Low coherence reflectometry for the characterization of fiber and planar

lightwave circuits," Symposium on Optical Fiber Measurements, Boulder, Colorado,

NIST publication 839, p. 159, (1992).

16. K. Takada, A. Himeno and K. Yukimatsu, "Jagged appearance of Rayleigh-backscatter

signal in ultrahigh-resolution optical time-domain reflectometry based on low-coherence

interference," Opt. Lett., vol. 16, p. 1433 (1991).

17. J. M. Schmitt, A. Knuttel and R. F. Bonner, "Measurement of optical properties of

biological tissues by low-coherence reflectometry," Applied Optics, vol. 32, p. 6032

(1993).

18. D. Huang, E. A. Swanson, C. P. Lin, J. S. Schuman, W. G. Stinson, W. Chang, M. R. Hee,

T. Flotte, K. Gregory, C. A. Puliafito, J. G. Fujimoto, "Optical Coherence Tomography,"

Science, vol. 254, p. 1 178 (1991).

19. F. P. Kapron, B. P. Adams, E. A. Thomas and J. W. Peters, "Fiber-optic reflection

measurements using OCWR and OTDR techniques," J. Lightwave Technol., vol. 7, p. 1234, (1989).

20. P. Lambelet, P. Y. Fonjallaz, H. G. Limberger, R. P. Salathe, Ch. Zimmer and H. H.

Gilgen, "Bragg grating characterization by optical low-coherence reflectometry," Photon.

Tech. Lett., vol. 5, p. 565, (1993).

21. P. Lambelet, R. P. Salathe, M. H. Garrett and D. Rytz, "Characterization of a

photorefractive phase conjugator by optical low-coherence reflectometry," Appl. Phys.

Lett., vol. 64, p. 1079,(1994).

22. E. Brinkmeyer and U. Glombitza, "Complex coherence-domain reflectometry in active

laser diodes," Opt. Lett., vol. 17, p. 1441, (1992).

23. J. F. Lucas, L. F. DeChiaro, C. Salla and C. Y. Boisrobert, "Low coherence reflectometry

and spectral analysis for detection of gain anomalies in semiconductor lasers," Electron.

Lett., vol. 28, p. 2085,(1992).

24. R. K. Hickemell, K. Takada, M. Yamada, M. Shimizu and M. Horiguchi, "Chromatic

dispersion of erbium-doped fiber amplifier measured by Fourier transform spectroscopy,"

Symposium on Optical Fiber Measurements, Boulder, Colorado, NIST publication 839, p. 201, (1992).

25. D. M. Baney and W. V. Sorin, "Extended-range optical low-coherence reflectometry

using a recirculating delay technique," Photon. Tech. Lett., vol. 5, p. 1 109, (1993).

95

POLARIZATION-MAINTAINING FIBER ALIGNMENT USING TUNABLE

POLARIZATION-MAINTAINING FIBER FABRY-PEROT INTERFEROMETERS

Yufei Bao, Kevin Hsu, and Calvin M. Miller

Micron Optics, Inc., 2801 BufordHwy, Suite 140, Atlanta, GA 30329, USA

Abstract

A unique and sensitive technique for aligning and characterizing polarization-maintaining fiber joints utilizing a

high-finesse polarization-maintaining fiber Fabry-Perot interferometer (PM-FFPI) and polarization resonance

splitting effect is reported. The alignment criterion, derived from the phase sensitivity of polarization extinction

ratio for elliptical inputs, is also applicable to general PM fiber splicing. PM fiber splices with high polarization

extinction ratio (> 38 dB) were fabricated using this method.

I. Introduction

Polarization-maintaining (PM) fibers have been used for various device applications in fiber optic

coinmunication systems [1-4], polarmetric fiber sensors [5,6] and laser systems [7], as well as input/output

pigtails in laser and integrated optical devices. In a PM system, birefringent axis alignment between

PM fibers is important since it directly affects the system performance. In this paper, we demonstrate a

unique and sensitive technique which utilizes the polarization resonance splitting effect inside a tunable PM

fiber Febry-Perot interferometer (PM-FFPI) to align and characterize PM fiber joints with linear or

elliptical inputs.

A PM-FFPI consists of a PM fiber cavity enclosed by two highly reflective plane mirrors. Due to

different mode indices associated with the two orthogonal polarization states inside a PM-FFPI, there are generally two distinct sets of resonances produced when cavity length is scanned, as shown on the

oscilloscope of the alignment system in Figure 1 . Cavity length tuning is achieved by stacked piezoelectric actuators. One set of resonance can be obtained only when a linear input polarization is aligned with one of

the birefringent axes of the cavity. Output polarization extinction ratio (PER) can be calculated from the

peak intensity ratio of the two polarization resonances. Therefore, a PM-FFPI can function as a

polarization analyzer, and be used for aligning and characterizing PM fiber joints.

II. PM fiber joint alignment

Figure 1 illustrates an example in which a PM-FFPI is used to align PM fibers or two PM devices. Two

alignment steps are involved here: (1) alignment between the birefringent axes of a PM fiber (PMF-1) and

the PM-FFPI at joint-2 as shown in Figure 1(a), and (2) alignment between the birefringent axes of PMF-1

and the second PM fiber (PMF-2) at joint-3 as shown in Figure 1(b). Alignment procedures for both steps are essentially the same. Alignment criterion is derived from the output PER and its maximum variation

range. PER is defined as:

PER = 10 log

ErtvE^.

10 oy^oy

(1)

97
where E^^, E^^ are the electric fields of the two polarizations after joint-2. For the first alignment step, E^

and E^y can be calculated by the following Jones matrix equation:

'-'OX = A

cos 02 sin 02

'-'oy -sin 02 cos 02 0 e

COS01 sin0i

-sin 01 cos 01

e"^°'^cos0o

e-"t'o/2sin0o (2)

where 0o and (|)q are input ellipticity and initial polarization phase difference of a polarized source, 0, is the

misalignment angle between input polarization and PMF-1 at joint-1, 02 is the misalignment angle between

PMF-1 and the PM-FFPI at joint-2, (j), is the polarization phase difference in PMF-1, and A is a complex

constant. PER can be a negative number if the ratio of E^yE*^y/E^j^E*^^ is less than 1 . It is easy to see that PER is a fiinction of ({), if 02 ^ 0. For example, for an input with 0^ = 0.5° and (()o= 10°, Figure 2 plots the

output PER versus (|), for different misalignment angle 02. Calculation results show that PER is constant

or independent of (j), only when 02 = 0, and that the sign of 02 causes a 180 degree phase shift in PER. This

constant PER (CPER) is determined by 0q, (J),,, and 0,. There are two positions of 02, symmetric to 02 - 0,

where the phase sensitivity to ^, is infinite. These two special positions of are important in aligning PM

fibers, and can be determined by the level of the CPER from the following equation:

02, = -023 = tan-' (10TM°). (3)

As shown in Figure 2, the CPER of -41 dB results in 02, = 0.51°. When |02| < 02,, 02 can be calculated

from the maximum variation of PER induced by (|),, and the value of CPER. Figure 3 shows the maximum

variation in PER as a fiinction of misalignment angles 02 for different CPER levels. When |02| < 02,, a

higher CPER level results in a higher PER phase sensitivity which enables a better alignment. For

instance, a maximum variation of 4 dB in PER indicates 10,1 ~ 0.12° if CPER = 40 dB , but represents |02l ~ 0.39° if CPER = 30 dB. For a fixed elliptical input, i.e. 0o and (|)oare fixed, the value of CPER is determined by misalignment angle 0,. Figure 4 plots the value of CPER as a fiinction of 0, for a 1° input

ellipticity and different Calculations show that the higher the input linearity (linear case: (^^ = nn, n = 0, 1, 2, ...), the higher the CPER that can be achieved by adjusting 0,. CPER is independent of only

when joint-1 is aligned, i.e. 0, = 0.

For a fixed elliptical input, theoretical calculations have shown the necessary condition for a proper

alignment ofjoint-2, which is that the PM-FFPI has a constant output PER while (j), in PMF-1 is changed.

The polarization phase difference in a PM fiber can be changed by either wavelength scanning or thermal

disturbance. Based on above calculation results, the procedure of aligning 02 is designed as follows: (1)

iteratively adjust 0, and 02 to obtain a maximum output PER, (2) with 0, fixed, adjust 02 while thermally

disturbing PMF-1, to locate the two maximum phase sensitive positions 02, and 02b, (3) rotate 02 to the center of 0,, and 02b, (4) with 02 fixed, adjust 0, to obtain the highest PER, (5) heat or cool PMF-1 and

monitor the maximum variation range of PER, (6) calculate 02 from PER^^^ and PER^^^, (7) rotate 02 in one direction by the calculated value and test the phase sensitivity, (8) if PER becomes a constant with

respect to (j), then the alignment is accomplished. However, if PER results in a higher phase sensitivity, then

rotate 0, to the other direction by twice the calculated value. Theoretically, a very accurate alignment can be achieved by repeating (6)-(8). One important issue should be addressed here is that the PM fiber must be

free of external stress in order to achieve accurate alignment. We have noticed that both PANDA and

elliptical-core fibers are extremely sensitive to random stress caused by epoxy used to bond PM fibers

inside glass ferrules [8], and that the external stress can completely degrade the alignment. Once 02 is aligned, the alignment of joint-3 follows the same process, except that now the polarization phase

difference in PMF-2 (instead of PMF-1) should be changed to examine the variation range of output PER.

98
II. PM fiber joint characterization

A PM-FFPI can also be used to characterize a previously spliced PM fiber joints. The characterization

procedure follows the alignment principles discussed previously. First, PMF-1 and the PM-FFPI at joint-2 are aligned by monitoring ^^ in PMF-1. Second, the angular alignment accuracy at joint-3 is calculated

using Eqs. (1) and (2) or estimated using Fig. 3 fi^om the measured PER^ and PER^^ while ^2 is changed.

The technique described above has been used to align and characterize several PM fiber joints. Table 1

lists the observed PER^ and PER^, calculated CPER, calculated alignment accuracy AG, and the

calculated output PER for a linearly polarized input (PERJ. It shows that all the PM-fiber joints aligned

with a PM-FFPI have high polarization extinction ratio for linear input (PERl> 34 dB). In order to achieve an accurate alignment or characterization, a detector with a wide dynamic range and a scope with high

resolution are required.

Table 1 No. PER_(dB) PER.. (dB) CPER (dB) A0(deg.) PER^CdB)

1 28 35 30.7 0.65 39

2 29 >50 35.2 1.12 34

3 28.7 33.8 30.9 0.45 42

4 26.4 27.2 26.7 0.12 53

5 29.7 32.6 30.9 0.30 46

6 25.5 30.9 27.7 0.70 38

rV. Summary

The alignment technique using a PM-FFPI is suitable for a PM fiber system which requires a PM-FFPI

as a wavelength or a noise filter. The birefringent axis alignment of the system can be easily accomplished

in one step by scanning the PM-FFPI. This approach requires no lens and analyzer, thus results in a simple

and convenient alignment process. PM-FFPIs can also be used in the following applications: aligning PM

fiber pigtails to polarization sensitive integrated optical circuits, and splicing additional PM fiber sections onto an existing PM fiber system.

In conclusion, we have developed a simple and unique technique to align and characterize PM fiber

joints using tunable PM-FFPIs. In addition to applications in wavelength and polarization scanning and

filtering, tunable PM-FFPIs are also sensitive instruments for aligning and characterizing PM fiber joints.

References:

[1] M. Asobe, H. Itoh, and K. Kubodera, Opt. Commun., vol. 88, no. 4-6, pp.446-50, 1992

[2] T. Morioka and M. Saruwatari, IEEE J. on Select. Areas in Commun., vol. 6, pp. 1186-98, 1988

[3] M. Kobayashi, K. takada, and J. Noda, J. I/g/i/w. Tec/zwo/., nol. 8, no. 11, pp. 1697-702,1990

[4] B.G. Clay, R. G. Podgomik, J. A. Priest, and C. L. Balestra, Conf. Digest. LEOS Summer Topical on

Optical Multiple Access Networks, p. 30-1, 1990

[5] S.-M. Tseng and C.-L. Chen, IEEE J. Lightw. Technol., vol!, no. 12, 1989

99
[6] Feng Zhang and John W. Y. Lit, Appl. Opt. vol. 32, no. 13, pp. 22U-1, 1993.

[7] I. N. Duling III and R. D. Esman, Electron. Letters, vol. 28, no. 12, pp. 1126-7, 1992

[8] Y. Bao, K. Hsu, and C. M. Miller, Technic. Digest OFC 94, pp. 96-8, 1994

Polarized source

thermal

disturbance

Joint- V 1 ^ Joint-2 i_Q_4 Tunable PM FFPI

PMF-1

Oscilloscope

(a)

thermal

disturbance V

Polarized source

Joint- 1 ^ Joint-3 ^ Joint-2 0 iJl4 Tunable PM FFPI

Scope

PMF-2 PMF-I

(b)

0.2 0.4

02 (degree)

Figure 1. PM fiber alignment illustration, (a) alignment of PMF-1 and PM-FFPI at joint-2, and

(b) alignment of PMF-1 and PMF-2 at joint-3.

Figure 2 PER versus phase retardation (j), for an

input with 1° ellipticity and 10° initial polarization

phase difference when 0,= 0.5°.

-10

-200 0 200 400

^ (degree)

600

Figure 3. Maximum variation in PER as a function

of misalignment angle for different output

constant PERs (CPER).

(degree)

Figure 4. For an input with 1° ellipticity and

different initial polarization phase retardations, the constant PER (CPER) is determined by 0, at joint- 1.

ICQ
Tapered fiber diameter measurements

Patrick Nebout, Nicolas Godbout, Suzanne Lacroix,

Xavier Daxhelet, and Jacques Bures

Departement de genie physique, Ecole Polytechnique de Montreal

P.O. Box 6079, Station •' Centre-ville" , Montreal (Quebec), H3C 3A7, Canada

1 Introduction

It is possible to locally reduce the diameter of a fiber by stretching it while heating it with a small flame. The resulting biconical structure, usually made from a single-mode fiber, is referred to a tapered fiber or, in short, a "taper" [1]. Fabrication parameters are important for they determine the longitudinal profile of the

structure which in turn determines the taper response, i.e., its behavior as a function of external parameters

such as wavelength, external index or incident power. Relatively large diameters, i.e., larger than 10 /zm, can be evaluated under microscope but the measurement precision is degraded as the diameter decreases. For diameters smaller than 10 /xm which are required for some appHcations, we developed a non-destructive

in-line measurement based on diffraction. Fiber tapers are made using the fusion tapering technique. A microtorch heats the fiber locally while the

fiber is pulled. The fabrication parameters like the pulling rate and the torch position must be controlled in

real time. For nonadiabatic tapers, i.e., abrupt tapers such as that of Fig. 1, the radius of the section being

pulled can be deduced from the oscillatory transmission of the taper with respect to elongation. However, in adiabatic tapers, the transmission remains by definition constant. Other methods are available but they are unpractical because they must be adapted to each particular taper. Since adiabatic tapers are required for certain appUcations, we need an alternate method that can ideally be used during fabrication. Because of

the small dimensions involved, a method taking advantage of the diffraction is appealing. It can also provide

the complete taper profile.

Section 2 is devoted to the theoretical analj'sis while Section 3 describes the experimental set-up and gives

the preliminary results. The last section dicusses prospects and draws the conclusions.

2 Theory

The problem of the transmission of a transparent fiber of radius a is a standard scattering problem by a two-dimensional body. The basis for the theory is well known, and can be found in many textbooks (see, for example, Ref. [3]). For the sake of simplicity, the fiber core effect is neglected and the observation is made

at an infinite distance. Let the incident monochromatic plane wave having a wavevector ko perpendicular to the fiber axis de- scribed in a polar coordinates system by

E'{r,9) = exp(fcor cos 61) (1)

For an incident electric field parallel to the fiber axis, the scattered field is shown to be

E'{r,e) = exp{ikor) exp(-z7r/4) j 6§ + 2 fe! cos(p0) V -^kor y

(2)

The series coefficients are given by _

kiJ!p{kia)Jp{koa) - koJp(kia)J^{koa)

^ ~

koJp{kia)Hi>'^'{koa) - kiH^'\kQa)J;{kia)

101
80

60 ^ 40

g

20

I

°

^ -20

£ -40

-60

-80

0 2000 4000 6000 8000

z (urn)

Figure 1: Profile of an abruptly tapered fiber. The untapered fiber parameters are: cladding diameter

(f)ci = 127 fj,m and core diameter (pco = 4.5 ^m; second mode cutoff wavelength Ac = 578 nm.

where ki is the wavenumber in the fiber; Jp and Hp^^ are the usual Bessel function and Hankel function of

the first kind of order p, and Hp^^' their derivatives.

The Fourier series of Eq. 2 was computed by adding the terms until their amplitude got smaller than

a convergence threshold e, typically e = 10"^. The computation of the series was done in double precision to avoid cancellation errors since a large number of terms is needed. To satisfy the convergence criterion,

the calculation for a 10 //m-diameter fiber requires 70 terms and a 125 /xm-diameter fiber, 750 terms. The

corresponding scattered field intensity profile is then deduced and compared to the experimental one.

3 Experimental results

The Fraunhoffer diffraction pattern is observed with the experimental setup schematized in Fig. 2. To test the

method, we made a measurement on a cylindrical fiber with a nominal diameter of 125 /xm. The experimental

recording using a He-Ne laser is shown in Fig. 3a. It is to be compared with the theoretical prediction in Fig. 3b where the diameter had to be adjusted to 127 /xm to obtain a fit in the position and relative amplitude

of the peaks. Preliminary measurements have also been made on tapered fibers. As an example, we show the

results for a tapered fiber with a quasi-cylindrical central region. We made a nonadiabatic taper and used

its oscillatory transmission to measure its diameter: 13.9 /Ltm. The experimental and theoretical diffraction

patterns are shown in Fig. 4. The position and relative amplitude of the first maxima and minima is very

sensitive to the diameter of the fiber which had to be adjusted to 13.74 ;um to obtain the agreement with a measurement precision of the order of 0.1%.

Apart from the diffraction spreading perpendicular to the fiber, an interesting effect has been observed

when the beam is pointed on the slopes of the taper. Two systems of interference fringes are then seen: the

usual diffraction pattern and another parallel one caused by the prism effect of the conical section of the fiber. Although we did not interpret completely these observations, it appears that the deflection of the secondary

fringe pattern is related to the taper slope. The measurement of the corresponding angle will then provide

additional information on the local diameter variation.

The results presented here are preliminary and a lot of experimental work remains to be done. The

method still has to be tested for very small diameters (under 10 /xm), for which it is most needed. For these

diameters, we lack an independent measurement to validate the method and thus additional information must

be drawn from other diffracted pattern measurements. Since the relation between the diffracted pattern and

the fiber dimension is not as simple as it would be for a slit, complementary calculations are also needed.

102
Laser

Figure 2: Setup used for the measurement. The parallel polarized laser beam is diffracted by the tapered fiber in the figure plane. A sUt perperdicular to the fiber isolates a region of constant diameter. The Fraunhoffer

diffraction pattern is observed in the focal plane of a convergent lens. The light is collected by a photodiode

through a multimode fiber which explores the diffraction pattern.

Figure 3: (a) Experimental recording of the diffraction pattern by a fiber with a nominal diameter of 125 /xm.

The wavelength was here 633 nm. (b) Calculated intensity pattern for a diameter of 127 yum.

103
10 I \ I I I I I I I \ I I i-j I 1—J I III I I I I 0 0.05 0.1 0.15

e

Figure 4: Experimental recording of the diffraction pattern by the central region of a tapered fiber having

a diameter of 13.9/xm (solid line). The dashed line shows the calculated intensity pattern for a diameter

adjusted to 13.74 ^m to obtain the best fit over the limited displacement of the detection fiber.

Different wavelengths might be used as well as different polarizations, the theory remaining to be extended

to the perpendicular case. The prism effect mentioned above must also be interpreted and exploited.

4 Conclusion

To measure the geometrical profile of tapered fibers, we developed a method based on the Fraunhoffer

diffraction pattern measurement. Despite the apparent complexity of the diffraction patterns, this method is promising because it is a nondestructive method, not depending of the taper adiabaticity and applicable for any diameter of interest ranging from 1 ^um to 125 /um. It still has to be refined by complementary theoretical as well as experimental work. In particular, the precision of the method must be explored systematically for different fiber diameters. In the longer term, it is expected to be reliable enough to be implemented in the

tapering setup for characterization of the longitudinal profile during taper fabrication.

References

[1] S. Lacroix, "Fused bitapered fiber devices for telecommunication and sensing systems," in SPIE 1994

International Symposium on Optics, Imaging, and Instrumentation, San Diego, CA, 1994.

[2] J. D. Love, W. .1. Stewart, W. M. Henry, R. J. Black, S. Lacroix, and F. Gonthier, "Tapered singlemode fibres and devices: Part 1. Adiabaticity criteria," lEE Proc. Pt.J: Optoelectronics, vol. 138 (5), pp. 343-354, 1991.

[3] H. C. van de Hulst, Light scattering by small particles, New York: Wiley, 1957.

104
Characterization of Silicon Optical Bench Waveguides

C. H. Henry and H. H. Yaffe'

AT&T Bell Laboratories, Murray Hill, N.J. 07974

Silicon optical bench (SiOB) waveguides are formed from doped silica films deposited on silicon substrates primarily by low pressure chemical vapor deposition. A typical

waveguide has a core 5x5 microns in cross section that is surrounded by 15 micron thick

base and cladding layers. The core -cladding refractive index difference is about 0.65 %.

These waveguides are well matched to standard optical fibers and have low loss per unit

length.

Optical integrated circuits (OICs) differ from fused fiber components in that complex

OIC waveguide paths can be precisely patterned by photo-lithography and etching with

excellent control over the phase differences between different paths of propagation.

Because of the large bend radii that are required for low loss, integrated optical circuits

tend to be quite long. Consequently, the large area substrates available in silicon

technology are advantageous. We use 5 inch diameter wafers at present. SiOB

components include couplers, taps, splitters, star couplers, broad and narrow band

multiplexers and low speed switches.

Waveguide insertion loss is a measure of the added loss of placing a waveguide between

a pair of optical fibers. This loss is mainly due to fiber-waveguide mode mismatch at the

two interfaces. The fibers are aligned to the waveguide with 3 axis piezoelectric

positioners. Matching oil is used to remove reflection loss and the effects of saw damage

at the facets. Adar and Serbin [1] have measured insertion losses at 1.55 microns of 0.5

dB for standard AT&T fiber and 0.1-0.25 dB using optimally matched fibers.

The loss per unit length is difficult to evaluate by direct measurements because the

attenuation length is much greater than the length of a straight optical path. We have overcome this by deducing losses from the finness of a ring resonator coupled to a

1 . Present address: AT&T Bell Laboratories, Norcross GA 3007

1

105
straight waveguide. The recent measurements of Adar et. al. [2] used rings up to 6 cm in

diameter. This measurement was aided by the use of an optical fiber laser which was

narrow in linewidth and could be thermally tuned without mode hopping [3]. Data on the

width of the resonance and the depth of the transmission dip at resonance can be used to separately evaluate both the resonator coupling loss and the loss per unit length. Losses

of less than one dB per meter were measured at 1.55 microns. The loss drops off slowly

with bend radius and the limiting loss mechanism is not known. The loss versus radius is shown in Fig. 1

An important application of this technology is in making multiplexers for

multiwavelength transmission systems with closely spaced channels. Such future

transmission systems are likely to have standard channel wavelengths and it will be

necessary to make optical components with transmission band centers within about 1 angstrom of these wavelengths. To determine what the wavelength control of our current

technology, we fabricated identical Bragg reflectors over a batch of wafers using deep

UV photo-lithography. Fig. 2 shows the distribution of measured Bragg wavelengths.

The distribution, which includes the data on 6 wafers, is about 3 angstrom wide. While

this spread is narrow, it probably can be improved. The wavelength variations are small

because silica has a stable refractive index and our waveguide effective refractive index

deviates only slightly from that of pure silica. The observed wavelength spread

corresponds to a 10 percent variation in core thickness, which is about what is measured.

A growing number optical waveguide components achieve spectral filtering by

controlling the relative phase of diverse paths from input to output. Adar et al. [4]

studied sets of nominally identical Mach-Zehnder interferometers (MZIs) to learn to what

extent we can control the phase in a waveguide path. Wavelength fluctuations of the

MZIs, Fig. 3, result primarily from process induced variations in the waveguide, such as

the roughness during etching. These induce a random walk in phase during propagation

down the waveguide that is characterized by a rms phase deviation is one radian at a

distance of 13 meters. While this is a long coherence length, these fluctuations have

noticeable effects that limit the performance of low order MZIs. Additional fluctuations were associated with the photo-lithographic masks. In this case, the pattern of

fluctuations repeats from wafer to wafer.

Strain plays an important role in determining the polarization dependence of optical

waveguides. The doped silica films are flowed during annealing at about 1000 degreeC.

106
Strain is relieved at this point. Upon cooling to room temperature, strain develops

because the silicon substrate contracts more with coohng than the doped silica films. The

compressive strain results in small polarization dependent refractive index changes in the

various doped silica materials. These changes produce a variety of effects including

small TE-TM splittings of narrow band filters. These splittings and the other polarization

dependent effects can be overcome by more sophisticated designs and additional

processing. Fig. 4 shows a narrow band Mach-Zehnder interferometer in which these

splittings are completely compensated for by the addition of a thin silicon nitride film in

part of one arm. [5]

Commercial application of OICs require nearly perfect optical waveguides with very low

loss, polarization independence, control of the absolute refractive index and control over

the optical phase difference between different paths. The SiOB waveguide technology

has the potential to meet all of these needs.

REFERENCES

[1] R. Adar, M. R. Serbin, and M. A. Meixler, "Reduced fiber coupling propagation

loss with modified phosphorus doped silica on silicon waveguides", Photonics

Technology Letters, (1994).

[2] R. Adar, M. R. Serbin and V. Mizrahi, "Less than 1 dB per meter propagation loss

of silica waveguides measured using a ring resonator", J. Lightwave Technol. (in

press) (1994).

[3] V. Mizrahi, D. J. DiGiovanni, R. M. Atkins, Y-K Park and J-M. O. Delavaoux,

"Stable single-mode erbium fiber-grating laser for digital communication", J. Lightwave Technol. vol 11, 2021-2025 (1993).

[4] R. Adar, C. H. Henry, M. A. Milbrodt and R. C. Kistler, "Phase coherence of

optical waveguides", J. Lightwave Technol., vol. 11, 212-219,(1993).

[5] H. H. Yaffe, C. H. Henry, R. F. Kazarinov and M. A. Milbrodt, "Polarizationindependent silica-on-silica Mach-Zehnder interferometers", J. Lightwave

Technol. vol. 12, 64-67, (1994).

107

Diagnosis of Silica-Based Waveguides by Using a

Jaggedness-Free OLCR

Kazumasa Takada, Hiroaki Yamada and Masaharu Horiguchi

NTT Opto-electronics Laboratories

Nippon Telegraph and Telephone Corporation

Tokai, Ibaraki-ken, 319-11, Japan

1.Introduction

The optical low coherence reflectometer (OLCR) is sufficientiy sensitive to detect

Rayleigh backscattering with a dynamic range of more than 20 dB [1-3]. Because of its high dynamic range and the high spatial resolution, the OLCR has great potential for use in

measuring the loss distribution in planar waveguides by using Rayleigh backscattering. The

measurement principle is the same as that by which OTDRs are used to diagnose optical fiber

cables. However, the OLCR has not yet been very successful when used for this

measurement [4-6]. This is because the Rayleigh backscatter signal jags randomly along its length due to speckle-like phenomena [7] so that the loss information is concealed by the

jaggedness.

This paper describes the loss distribution measurement of silica-based waveguides

using a jaggedness-free OLCR. A tunable low-coherence light source is specially designed for

the efficient wavelength averaging [8] of the Rayleigh backscatter signal. The OLCR with this

averaging scheme, or jaggedness-free OLCR, reduces the jaggedness sufficiently to reveal

the mean Rayleigh backscattering profiles of the silica-based waveguides. Their loss

distributions can be clearly determined from the profiles.

2.Experimental Setup

2.1. OLCR Svstem : Figure 1 shows the configuration of the OLCR with a [3x3] fiber

coupler. The output from the low-coherence light source, which is shown schematically in

Fig. 2, is launched into the [3x3] coupler from the source port, and is divided equally into

three ports on the output side. Two are used as the probe and local oscillator (LO) ports,

while the other is unused. The piezoelectric (PZT) phase modulator incorporated into the

probe port is sawtooth ramped through a k radian range in order to produce a serrodyne

carrier frequency [9]. The polarization controller (PC) is used to excite one principal mode of

the waveguide under test (WUT). The light from the LO port is emitted into the air and

reenters the port after being reflected twice at the movable retroreflector and once at the mirror.

The LO light and the light returned from the WUT are combined by the coupler. The combined

light is detected by a pair of photodiodes after propagating through the two balanceddetection ports on the input side of the coupler.

2.2. Low-Coherence Light Source : The previously-reported tunable low-coherence light source [8] has the following disadvantages. The output is limited to a few milliwatts. TTie source can be considered to be a Fabry-Perot resonator, in which a fiber amplifier is sandwiched between a mirror and an optical isolator. Since the return loss of the isolator

is no more than 60 dB, the round-trip gain in the cavity at higher pump powers exceeds the

round-trip loss of about 60 dB. Even when the laser oscillation is suppressed, the gain near

the loss produces undesired satellites in the Fresnel response function of the OLCR,

widening the hidden space near the facets and an intense WUT reflection.

A tunable high-power low-coherence light source was constructed as shown in Fig.

2. This source uses two optical amplifiers AMPl and AMP2, which are optically

interconnected in such a manner that the laser oscillation is suppressed by optical

isolators ISOl and IS02. Each amplifier was composed of an erbium-doped fiber, a WDM

coupler, an optical isolator and a 1.48-)im pump LD. The optical bandpass filter module

FILl was used for filtering the broadband superfluorescent light from AMPl. The narrow- band hght from the filter, which is amplified by AMP2, is passed through the second bandpass

109
filter module FIL2 in order to remove the wideband spontaneous emission generated by

AMP2. These bandpass filters are automatically tuned by a wavelength controller so

that the center wavelengths of the passbands coincide.

One end of the erbium-doped fiber of AMPl is polished at an angle of 8°, thereby

giving a return loss of 60 dB. The total return loss of ISOl and FILl is 60 dB. The light

suffers a total loss of approximately 120 dB during a round trip in the Fabry-Perot resonator

which is set up between ISOl and the angled-endface of the erbium-doped fiber. Since the

loss greatly exceeds the gain of AMPl, the optical feedback effect can be greatly reduced.

Since the second amplifier AMP2 is sandwiched between ISOl and IS02 whose return

losses are 60 dB, respectively, the effect of the feedback can be reduced as in AMPl.

The center wavelength of the source output was tuned from 1560 nm to 1530 nm

in approximately 1 nm steps. Figure 3 shows the optical spectra of the output on a

logarithmic scale. The pump power launched into the optical amplifiers was 60 mW. The

rotation angle in the figure is that of the bulk-optic filter incorporated into the second

bandpass filter FIL2. The output power from the source changed from 10 dBm at 1560 nm

to 6 dBm at 1530 nm. Although the power from the source was very high, each output

was narrow-band, low-coherence light whose bandwidth was less than 1 nm and whose

spectrum had sidelobes that were -40 dB lower than the central peak. We launched each of

these low-coherence hghts into the OLCR system and found that a deviation of 2.5 mm from a

silica-based waveguide endreflector resulted in a signal that was 60 dB lower than that detected

at the reflector. Thus the hidden space defined by the full width at -60 dB maximum of the

Fresnel response is 2.5 mm x 2= 5 mm,

2.3. Wavelength Averaging : The retroreflector of the OLCR was scanned at 250 [im

intervals, and reflection from silica-based waveguides was detected at each position of the

retroreflector. The mean value of the Rayleigh backscatter signal detected by the OLCR is proportional to P^/ 5v, where P is the source output, and 5v is its effective bandwidth [10].

N

The wavelength averaging at each position was carried out by calculating iy^CPy/'J2c 5^^^/

d\)s)IN after sweeping the entire wavelength range once from Xi =1560 nm to A^/ =1530 nm

{N = 30) in approximately 1 nm steps. Here the monitored source output P)^ and the

detected signal were stored in a computer for each wavelength A^. 5vk is the effective

bandwidth at A^. Ps and 5vs are the power and the effective bandwidth of the output at a standard wavelength of Ay = 1555 nm.

3. Results and Discussion

First, a high loss silica-based waveguide fabricated on a silicon substrate was

tested. The waveguide was connected via index matching oil to single mode fibers, one of

which was the fiber probe of the OLCR. The TE mode of the waveguide was excited at the input end. Figure 4 shows the mean reflection profile obtained after performing

wavelength averaging. The profile is composed of a "Rayleigh backscattering" floor that

is revealed by averaging the Rayleigh backscatter signal and from which the loss distribution

is obtained, isolated (or point-like) reflections distributed randomly along the waveguide,

and an endreflection at its input end (at zero distance). The fluctuations in the floor of the

probe fiber (distance < 0) were reduced to within ± 1 dB by using the wavelength

averaging method. The index matching oil reduced the endreflection to only -40 dB, whereas

the revealed floor level was less than -95 dB.

In spite of such a large endreflection, it was observed that the floor (measured in dB)

of the waveguide attenuated linearly with distance by 17 dB in 40 cm, and the attenuation

was not affected by the location and reflection of isolated scattering centers. The waveguide

receives a constant loss over the waveguide at a rate of (17 dB)/(40 cm)/2 = 0.2 dB/cm,

which agrees well with the directly measured value. The behavior of the floor suggests that

the origin of the waveguide loss is not these isolated scattering centers, but results from

fluctuations such as core-cladding irregularities or inhomogeneity in the core and cladding.

110
Next, we tested silica-based waveguides fabricated on a different substrate. Nineteen

waveguides were fabricated in parallel on the substrate in a "serpentine" configuration and

were numbered from 1 to 19 in sequence. With the conventional method for estimating total loss, two waveguides (Nos. 1 and 18) were proved to have extra losses of 4 dB and 8 dB,

respectively, as compared with the other waveguides. Figure 5 shows the mean reflection

profile of one normal waveguide (No. 7). Although this waveguide had large point-like

reflections, the Rayleigh floor attenuated linearly with distance by 4 dB in 40 cm. The

waveguide receives a constant loss of (4 dB/40 cm)/2 = 0.05 dB/cm along its length.

The mean reflection profiles of five silica-based waveguides (Nos. 1, 2, 3, 7, and 18) are shown in Fig. 6. The Rayleigh floors of the two normal waveguides (Nos. 2 and 3)

exhibited approximately the same attenuation behavior as that of waveguide No. 7. The

floor of the waveguide No. 1 decreased in steps of 2 dB and 5 dB at distances of 18 cm and 36

cm, respectively. Waveguide No. 18 decreased by 14 dB at a distance of 26 cm. These

reductions mean that extra losses of 2 dB/2 -i- 5 dB/2 = 4 dB, and 14 dB/2 = 7 dB are added

for waveguides Nos. 1 and 18, respectively, whose values agree well with the extra

directly measured losses. The loss coefficients at portions other than these points are approximately the same as those of the normal waveguides. These results indicate that the

measured extra losses are produced by point defects. One possible origin of the defects is the

unexpected adhesion of small particles of dust on the waveguides.

4. Conclusion

We have described the loss distribution measurement of silica-based waveguides by

using a jaggedness-free OLCR. A tunable low-coherence light source using two optical

fiber amplifiers was specially designed for the diagnosis. The source automatically tuned the

central wavelength of the low-coherence light with a spectral width of less than 1 nm over 30 nm. Unavoidable jagged fluctuations observed when using conventional OLCRs were

reduced to within + 1 dB by averaging the wavelength-dependent Rayleigh backscatter

signals. The source reduced the optical feedback effect sufficiently to limit the hidden

space defined by the full width at -60 dB maximum of the Fresnel response to 5 mm. This

OLCR with the wavelength averaging scheme, or jaggedness-free OLCR, revealed the

Rayleigh backscattering floors of the waveguides in spite of large endreflections at their

input ends. Constant loss distributions such as 0.05 dB/cm and stepwise loss increases of, for

example, 1 dB, were clearly observed in the waveguides from the changes in the floors.

5. Acknowledgments

The authors would like to thank Dr. Masao Kawachi for his continuous

encouragement. They also thank Dr. Yoshinori Hibino for supplying us with the waveguide

samples used in the experiment.

References

1. K. Takada, M. Shimizu, M. Yamada, M. Horiguchi, A. Himeno, and K. Yukimatsu,

Electron. Lett. vol. 28, pp. 29-30 (1992).

2. W. V. Sorin and D. M. Baney, IEEE Photon. Technol. Lett. vol. 4, pp. 374-376 (1992).

3. K. Takada, T. Kitagawa, M. Shimizu, and M. Horiguchi, Electron. Lett. vol. 29,

pp. 365-367 (1993).

4. K. Takada, N. Takato, and J. Noda, Opt. Lett. vol. 14, pp. 706-708 (1989).

5. M. Kobayashi, H. Hanafusa, K. Takada, and L Noda, IEEE J. Lightwave Technol.

vol. 9, pp. 623-628 (1991).

6. H. H. Gilgen, G. Bodmer, and Ch. Zimmer, in proceedings of 2nd Optical Fiber

Measurement Conference, Torino, Italy, pp. 143-146 (1993).

7. K. Takada, A. Himeno, and K. Yukimatsu, Opt. Lett. vol. 16, pp. 1433-1435 (1991).

8. K. Takada and M. Horiguchi, Electron. Lett. vol. 29, pp. 1273-1274 (1993).

9. R. P. Novak, H. H. Gilgen, P. Beaud, and W. Hodel, in Proceedings of Symposium

on Optical Fiber Measurement, Boulder, Colorado, pp. 35-38 (1990).

10. P. R. Morkel, R. I. Laming, and D. N. Payne, Electron. Lett. vol. 26, pp. 96-98 (1990).

Ill
analyzer

AMP1

erbium-doped

fiber

AMP2

FIL1 IS01

erbium-doped

fiber

^^^^

WDM -^-(^

pumpC^

IS02 FIL2

output

retroreflector

wavelength

controller

Fig. 1. Experimental setup of OLCR

PC: polarization controller, LO: local oscillator,

WUT: waveguide under test

Fig. 2. Experimental setup of tunable lowcoherence light source

FIL: optical bandpass filter, ISO: optical isolator,

AMP: optical amplifier

OQ>

>

I-OLU

UJ

1520 1530 1540 1550 1560

WAVELENGTH (nm)

1570

u 1 1 1 - I r

1 1 -20

3 probe 2 -40

1 waveguide

-60

-80

-100

Vlu.. .1 . L -120

-140 .. . 1 1 1 1 1 1 1

Fig. 3. Change in output spectrum from

the low-coherence light source

0 10 20 30 40

DISTANCE (cm)

Fig. 4. Mean reflection profile of a silicabased waveguide with high loss

20 dB

10 20

DISTANCE (cm

Fig. 5. Mean reflection profile of a silicabased waveguide

DISTANCE (cm)

Fig. 6. Mean reflection profiles of silicabased waveguides fabricated on

one wafer

112
MEASUREMENTS OF LOSS AND MIRROR REFLECTIVITY IN

SEMICONDUCTOR OPTICAL WAVEGUIDES: A EUROPEAN

INTERLABORATORY COMPARISON EXPERIMENT

E. Gird - Institute of Quantum Electronics, ETHZ, Zurich, Switzerland

L.H. Spiekman - Dept. of Electrical Engineering, TU Delft, Delft, The Netherlands

H. van Brug - Dept. of Applied Physics, TU Delft, Delft, The Netherlands

J.-F. Vinchant - Alcatel Alsthom Recherche, Marcoussis, France

S. Morasca, F. Pozzi, C. De Bemardi - CSELT S.p.A. , Torino, Italy R. Zengerle, W. Weiershausen, W. Noell - DBP Telekom, Research Center Darmstadt, Germany

L. Thevenaz, A. Kiing - EPFL, Metrology Laboratory, Lausanne, Switzerland

A. Enard, N. Vodjdani - Thomson CSF Laboratoire Central de Recherche, Orsay, France

The results of round-robin measurements of waveguide loss and integrated mirror

reflectivity in several European laboratories in the frame of the COST 240 project are presented.

The Fabry-Perot cavity contrast method was used on semiconductor optical waveguides. The

measured losses show a satisfactory agreement (within ±0.5 dB) between all the laboratories, but

also point out some aspects which must be taken into account to get meaningful data.

1. - Introduction

Several aspects of the measurement procedure for integrated optical waveguides and

components are at a far less evolved stage than for the optical fiber counterparts. The repeatability

and reproducibility of measurements, the homogeneity and comparability of results from

equivalent or similar methods or set-ups, even for basic parameters like waveguide loss, are just a few of these aspects. To address this problem, within the European Project COST 240 "Techniques

for Modelling and Measuring Advanced Photonic Telecommunication Components" it was decided to organize the circulation of waveguide samples in a number of participating laboratories, to perform

measurements and compare the results.

As a first test, it was agreed to measure the loss of waveguides fabricated on InGaAsP/InP

chips, using the well-known method of Fabry-Perot resonances contrast in the cavity formed by the

partially reflecting terminal facets [1]. The choice of this technique was prompted by its sensitivity, non destructiveness and availability (in a few modifications of the basic configuration)

at several participating laboratories.

2. - Principle of the technique and experimental set-ups

In a Fabry-Perot resonator operating in a single transverse mode regime, the ratio K

between its maximum and minimum transmittance when changing the resonance conditions appears

in the following relationship [1]:

where L is the geometrical length of the cavity (i.e. the waveguide in the present case), a the

cavity attenuation and R the mirror (facet) reflectivity, assumed equal for both ends.

The basic experimental set-up to use this technique for waveguide measurements is shown in

Fig. 1; the actual implementations differ in several details, the most significant one being the way

ABSTRACT

(1)

113
of changing the resonance conditions to scan several periods of the transmission characteristics, and

hence to determine K. Two techniques are currently in use: a) a slow heating or cooling of the

sample, by which the length of the optical cavity is changed by the combined action of thermal

expansion and thermooptic effect, while the wavelength of the narrow-linewidth source is kept

fixed; b) small changes of wavelength of the source, while keeping constant the sample

temperature and its optical length. AAR, DBF Telekom, EPFL and Thomson CSF used method a), while ETHZ, both TU Delft laboratories and CSELT used method b).

Other minor differences pertain to the source (gas or diode laser), the way of coupling light to the waveguide (lens or single-mode fiber) and the type of control of the input or output

polarization state. All the implementations in the participating laboratories use either HeNe

lasers operating at 1.532 fim or diode lasers operating in the 1.55 |j.m range.

3. - Sample structure

Semiconductor chips with stripe waveguides fabricated in the InGaAsP/lnP material

system were made available for this experiment; the high refractive index of these materials

yields natural Fresnel reflectivities for the facets around 0.3, so that good contrast can be achieved

without special sample preparation besides cleaving the I/O facets. A first sample with simple

straight ridge guides, supplied by CSELT, was used for some initial essays; then a more complex

chip, containing both straight guides and different types of integrated mirrors, supplied by ETHZ,

was circulated among all the labs and extensively characterized. Its structure is shown in Fig. 2, and all data reported in this paper refer to this sample.

4. - Measurement results and discussion

Measurements of Fabry-Perot contrast were performed on selected waveguides and mirrors on the chip, separately for TE and TM polarizations. From the measured values of K, the total cavity loss was calculated, and the corresponding values are shown in Fig. 3 for the TE

polarization. The data show a general agreement between the results of all laboratories, with some

occasional discrepancies and more scattered data for particular guides; in two cases (Thomson-CSF

and ETHZ data) the measurements were made after a new cleavage of one end (sample length

reduced from 9 to 5.6 mm), so that the facets are not the same as for the other laboratories, and

their quality may be different. Other differences can arise from cleaning of the sample in different

solvents, which was performed in some laboratories to remove the mounting wax used to fix the

chip on the holder after the measurements: residual contamination can alter both R (if the facet is affected) and a (if the guide surface is contaminated); a possible example of this effect is the

sequence of results from AAR, CSELT and DBP for guides on one side of the chip. Some guides 6 |im

wide exhibited multimode behavior, and in this case the contrast values K depended more or less sensitively on the input coupling conditions, thus increasing the uncertainty.

To get the corresponding values of attenuation a (expressed in dB/cm) for the waveguides,

in addition to the measured length L, calculated values of modal reflectivity were used for R; these were computed by a Fourier analysis and Fresnel reflection method [2] applied to the actual

waveguide structure, and range between 0.315 and 0.319 (TE) or 0.247 and 0.250 (TM) according to the

guide width; no experimental value for R was available.

The attenuation values of course exhibit the same trend as the total cavity loss, and the

average level is 1.1 ± 0.3 dB/cm for TE polarization; as a comparison, it should be noticed that for

this chip length an uncertainty of ±0.1 in the corresponding contrast (K=2.8) amounts to ±0.16

dB/cm, and increases to ±0.5 dB/cm when K=1.75. An uncertainty of ±0.01 in the adopted value of R

leads to a corresponding systematic error of ±0.16 dB/cm in the attenuation.

114
The results for TM polarization are similar, except for a larger spread, due to lower

reflectivity, smaller K values and higher loss (a = 2.4 ± 0.6 dB/cm).

The reflectivity of integrated mirrors was obtained by subtracting the average guide loss of

the corresponding straight guides; mirror losses of 0.8 ± 0.2 dB and 1.0 ± 0.3 dB were obtained for TE

and TM pxjlarization respectively.

5. - Condusi(xis

An extensive series of optical loss measurements has been performed by several laboratories on semiconductor waveguides by similar set-ups based on the Fabry-Perot fringe contrast method.

Agreement within about ±0.5 dB/cm has been found in the vast majority of cases for guide

attenuation, and within ±0.2 dB for integrated comer mirros. It has been shown, on the other hand,

that maintaining very carefully facet quality and cleanliness is extremely important, and that

definitely single mode propagation is essential to achieve meaningful and reproducible results.

References

[1] R.G. Walker, "Simple and accurate loss measurement technique for semiconductor optical

waveguides". Electron. Lett., vol. 21, pp. 581-583, 1985.

[2] P.A. Besse, J.S. Gu, H. Melchior, "Reflectivity minimization of semiconductor laser amplifiers

with coated and angled facets considering two-dimensional beam profiles", IEEE J.Quantum

Electron., vol. QE-27, pp. 1830-1836, 1991.

NLWS ICO

WG —TPTC

oco D

PM

Fig. 1 - Schematic of a basic set-up for measurements of Fabry-Perot fringe contrast in optical

waveguides. NLWS: narrow linewidth laser source with optical isolator; ICO: input coupling

optics; WG: waveguide under test; PTC: sample position and temperature control; OCO: output

coupling optics; D: detector; PM: power meter; CU: control and acquisition unit. Polarization control

is present at least in one coupling optics unit.

115
InP

1 /vm

InGaAsP 0.55

0.63 /ym

]

InP 1/;m

n-lnP /

Fig. 2 - Schematic structure of the layout and waveguide cross-section of the ETHZ-suppHed chip.

Groups 1 and 7 contain sets of straight waveguides 3, 4, 5 and 6 ^im wide; group 2 contains pairs of 45°

deviating mirrors; groups 3 to 6 contain pairs of 90° deviating mirrors; mirror waveguides have the same widths as straight guides; total chip length 9 mm. The facets have no intentional coating.

10

CD 8

O 7

>

CO O

A

C) o

1 /'\ V

i

A o

o

i

a

I

A

O

o o AO ^ A i S H

B

o o ^

• o o

A

A

A A

o

o

Wavequide ^uide W

5

Hdth [\xm]

Waveguides with

Corner Mirrors

(45° Deviation)

Waveguides with

Corner Mirrors

(90° Deviation)

Straight Waveguides

Fig. 3 - Distribution of the measured total cavity loss (propagation plus reflection loss at the facets) for TE polarization. O : TUD-EE; : TUD-AP; O : AAR; A : CSELT; • : DBF Telekom; : EPFL;

: Thomson-CSF; A : ETHZ (in order of measurement). The line at the 5.0 dB level corresponds to the calculated end facet reflection loss; average attenuation for waveguides is 1.1 ±0.3 dB/cm (see

text). The bars at left show the uncertainty coming from a AK = ±0.1 at different total loss levels.

116
White-Light Spectral Analysis of Lithium Niobate Waveguides

C. F. McConaghy, K. F. Hugenberg, D. Sweider, and M. Lowry

Lawrence Livermore National Laboratory

R. A. Becker

Integrated Optical Circuit Consultants

1. INTRODUCTION

Optical channel waveguides in lithium niobate are a very useful technique for building a variety of

integrated optical components such as modulators, splitters, and frequency doublers. In many of

these applications, single mode performance is essential in order to achieve the desired device

performance. Determination of single-mode operation can be complicated by the fact that a multimode guide can support lower order modes and can appear single mode if the launch conditions for

the hght into the guide do not excite any of the higher order modes. Typical measurements often

made to assure single-mode operation include optical mode spot imaging, and the "bhnk test". While these measurements are useful, they are stricdy quaHtative in nature, and thus do not provide

the fine detail often desired in practice. In developing waveguide fabrication processes, it is more

helpful to know how various process steps affect the wavelength at which the waveguide is singlemode, multi-mode and cut-off. In this paper, we show the usefulness of exciting waveguides with

a broadband white-light source and observing the output of the guides over much of the visible and

near IR spectrum with a spectrometer. Although this technique is not new, its usefulness in

process development and control has not been emphasized^.

2. EXPERIMENTAL SETUP

Figure 1 shows a block diagram of the white light spectrometer system. A tungsten-halogen lamp

is coupled into and out of the waveguide under test by microscope objectives. The lenses which

have a numerical aperture of 0.2 are achromatic and fairly well corrected in the visible and IR

portions of the spectrum. In addition, a calibration trace is done to calibrate out variations of the

lamp and optical system with wavelength. It should be noted that an ideal Gaussian spot when

focused will only excite the lowest order mode. However, there is some coupling to the next

highest order mode since the actual system is not a perfect Gaussian. The output tight is then

analyzed with a monochromator and a PIN diode detector. An ideal white light trace is shown in

Figure 2. Assuming that the wavelength spans over the range of the lowest order mode and the

next higher order mode, tiie ideal trace can be broken into four regions^. In Region 1 both modes

are far from cutoff and the ratio of light intensity at the output of the guide I(X,zo) to the total available input Ught intensity Is(A.) can be described as:

I(X,zo)/Is(A) = Koo(^) + Koi^) (1)

where X is the wavelength, zq is the length of the guide, and Kqo and Kqi are the input coupling

coefficients for the lowest and next higher order modes, respectively. Region 2 describes the X

regime where the Iqi cutoff is approached. In this case the ratio of output to input intensities is given by:

I(X,Zo)/Isa) = Koo(^) + Koi(^)exp(-aoi(?i)zo) (2)

where aoo and aoi are the modal attenuation coefficients for the lowest order and next higher order

modes, respectively. As tiie wavelength becomes longer the cutoff of the Iqi mode is passed and

region 3 covers tiie wavelength range where the Iqo mode is still far from cutoff. This intensity

ratio in this region is given as:

117
I(X,Zo)/Is(?^) - Koo(?i) (3)

Finally as the Iqo mode cutoff is approached in region 4, the intensity ratio is :

l(X,zo)fls('^) = Koo(^)exp(-aoo(?^)zo) (4)

3. EXPERIMENTAL DATA

For this experiment a mask with straight guide dimensions of 2,3,4, and 5 um was used to pattern

channels in oxide on three-inch diameter x-cut lithium niobate wafers. These wafers were then

proton exchanged in benzoic acid at 200°C and annealed at 350°C. Figure 3 shows a series of

experimental traces corresponding to five different wafers processed at different times over a 12

month time period. These particular guides had a mask width of 3 um and were exchanged for 10

minutes and annealed for 2 hours. The target wavelength for single mode operation of devices

made with these guides was 830 nm. It is evident from Figures 2 and 3 that 830 nm is at the

beginning of Region 3, just beyond the Region 2 edge for all five wafers. This wavelength

corresponds to the tightest confinement of the lowest order mode and occurs just prior to the

multimode region. The white light data provides a good measurement of the waveguide process

control. For example, the slight variation in wavelength of the regions for the various wafers can

be attributed to the degree of control in maintaining the mask linewidth in the masking oxide used on the wafer. Also shown in Figure 3 is a simulated trace assuming the following parameters:

ao(?c ) = 5.2 X 10-10 X (X-nOO) Nepers/nm ; Ko(>.) = -.0202 + 2.0 x 10-5 x X

ai(X) = 2.1 X 10-10 X (X-775) Nepers/nm ; Ki(X) = 0.0128 - 8.0 x 10-6 x X

where as previously mentioned, a simple linear dependence with wavelength is assumed.

Another interesting way to present the white light data is shown in Figure 4. Here, the

wavelengths for absolute single mode cut-off, well confined single mode operation, absolute

second mode cutoff and well confined double mode are plotted vs. mask width dimension. By

looking at the well confined single mode region , it is possible to select the proper mask width to use for a given wavelength of operation. It is apparent from this figure that the well confined single

mode regime covers a broad spectrum (350 nm for the 3 um guide). Experiments have also been

conducted with devices under bias to observe the lowest Vti wavelength on a Mach Zehnder and

the change in cut-off wavelength under bias on a phase modulator.

4. SUMMARY

We have shown the usefulness of exciting waveguides with a broad spectrum from a white-Ught

source and measuring the guide response vs. wavelength. By proper interpretation of the various

steps in these responses, it is possible to quantitatively determine cutoff, single mode operation,

and multi-mode regimes and to monitor process variability over time. In conclusion, the white

Ught excitation and spectral analysis technique is an important diagnostic for waveguide and device

development in lithium niobate and any other optical waveguide material system.

1. J.E. Baran, Y. Silberberg, P. Perlmutter, "White-light characterization of integrated-optic

waveguides and devices," OFC/IOOC, TU02, 1987.

2. K.A.H. van Leeuwen and H.T. Nijnuis, "Measurement of higher-order mode attenuation in

single-mode fibers: effective cutoff wavelength," Opt. Letters, 9, 252 (1984).

Work performed under the auspices of the U.S. Department of Energy by Lawrence Livermore

National Laboratory under Contract No. W-7405-Eng-48.

118
6.3X Obj. Piezo Driver

for Objective

Chofjper Driver Bye I

Tungsten

Halogen

Lamp

Piezo Driver

for Wafer Monochromator Driver

Fig. 1 Experimental setup for exciting waveguides with white light: includes a tungsten-halogen lamp, microscope objectives, apertures, monochromator,

and PIN diodes.

V3c

3

a.3

o

D

N

O

Regions

1 2 3 4 — — — — ^

KooC^a^+KoiC^tj)

^ ^ ^01 Cutoff %Cutoff \

Fig. 2 a) Typical data for a lithium niobate waveguide scanned from 600 to 1700 nm

b) Ideal data showing four regions:

(1) Iqo and Iqi modes ; (2) Iqi cutoff ; (3) Iqo mode ; (4) Igo cutoff

119
+

600 700 800 900 1000 1100 1200 1300 1500 1600 1700

Wavelength ( nm)

1400

Fig. 3 White-light spectrum for five non-simultaneously processed wafers with a target

single-mode wavelength of 800 nm. Also shown is a simulated trace.

Absolute Single

Mode Cut-Off

Well-Confined ^

Single Mode

Absolute Second

Mode Cut-Off

\

Well-Confined

Double Mode

2.5 3 3.5 4 4.5 Waveguide Mask Width (um)

Fig. 4 Wavelengths of operation for single mode,multi-mode, and cutoff vs. mask

dimensions for lithium niobate proton exchange guides fabricated with a 10 minute

200*C exchange and 2 hour 350*C anneal.

120
Frequency response measurements of integrated-optic electrodes

K. F. Hugenberg, P. D. Sargis, C. F. McConaghy

Lawrence Livermore National Laboratory

ABSTRACT

The frequency response of electro-optic waveguides can be determined using a variety of testing

methods. In this paper, we compare and contrast three measurement techniques used to test our LiNbOs devices for improving packages and electrode designs. Each method is described and

accompanied by typical results and the experimental setup. Finally, we summarize the advantages

and disadvantages of each method.

For the purpose of this paper, we use a Mach-Zehnder (MZ) device to demonstrate our measurement techniques for characterizing electro-optic devices. The operation of a Mach-Zehnder

device is worthy of review. 1 The phase difference between the two waveguides can be electrooptically induced to form the familiar interference modulation curve. The lowest order mode is excited when the two arms are in phase at their recombination point, and this produces a maximum

in intensity out of the device. The minimum of the curve arises when the two arms are 180 degrees

out of phase. The recombination in this case excites the first antisymmetric mode or second-order

mode which is cutoff. Consequently, the light radiates into the substrate and the transmitted light

out of the device is a minimum. The output intensity of the MZ is expressed by:

The function, m(f), represents the frequency dependence of the modulator. It includes the effects

of mismatch between the optical and electrical velocities (i.e., walkoff)!, the electrode's

conductance, and other factors related to packaging.2 Vj^ represents the half-wave voltage, and

0/2 is the static phase imbalance in the arms of the interferometer.

We use a traveling-wave^'^ titanium-indiffused, lithium niobate Mach-Zehnder to demonstrate our measurement techniques. The electrodes consist of a 2S-Q coplanar strip transmission line having

a strip width of 15 \im, a length of 3 mm, a gap width of 10 |J.m, and a gold-plating thickness of 4

|j.m. They are fabricated on a 0.2-|im oxide buffer layer to reduce metal loading losses. The rf input of the MZ device is interfaced to a 50- Wiltron K connector using a 4-mm long coplanar

taper transition. A 2S-Q. termination resistor is included in the package design for impedance

matching. The extinction ratio exceeds 20 dB, and the half-wave voltage (Vjj) is 27 volts. The

1300 nm single-mode optical waveguides are fabricated on x-cut, y-propagating LiNbOs

substrates.

It is well understood that the frequency response of a traveling- wave electrode design is limited by

the optical and electrical velocity mismatch (or "walkoff), and the frequency-dependent rf

propagation loss.^ We have found that knowledge of the frequency response, m(f) (see eqn. 1),

provides very useful information for improving our packaging and electrode designs. We describe

three techniques for measuring this parameter.

1. INTRODUCTION

(1)

2. FREOUENCY RESPONSE MEASUREMENTS

121
2. 1 Swept-frequency method

In Fig. 1, a swept-frequency synthesizer is amplitude-modulated by a constant low-frequency sine wave. The magnitude of the envelope from the amplitude modulation is then detected using a slow-response detector and a lock-in amplifier. The analysis shows that if the optical bias and the

driving source voltage are kept constant, then the rf frequency dependence of the low-frequency

signal amplitude directly corresponds to the modulator response.^

The optical bias was set at the peak intensity point on the electro-optic transfer function throughout

the measurement. The swept-frequency synthesizer was used to drive the rf input from 0.01 to 20

GHz. The rf power input to the modulator was 0 dBm and the repetition rate of the modulating

sine wave signal was 5 kHz. The amplitude was kept much less than V;,; to satisfy the small signal

requirements. (Results are in Fig. 6 (a).) When the driving voltage amplitude is kept small, then

the following applies to the optical intensity (Id) on the detector^:

lda[m(f)]2 ; m(f) refers to eqn. (1) (2)

2.2. Network analyzer method

The second measurement technique uses a network analyzer. Its function is to compare the

transmitted and received rf power as a function of frequency relative to a test or reference path, and

it results in a direct measurement of the small-signal frequency response.^

The experimental setup is shown in Fig. 2. The optical bias was set in the linear region of the

transfer function curve throughout the measurement. A 20 GHz InGaAs photodiode was used to convert the optical modulator output back to an electrical signal. The network analyzer was

calibrated to measure the forward transmission coefficient (S2i) through the device over a frequency range of 0.045-20 GHz. The rf power at the input of the modulator was approximately

-5 dBm. The effects of cables and connectors were eliminated in the calibration process by

employing standard techniques for measurements with a network analyzer.

The network analyzer measures frequency response by generating a swept cw signal, and

comparing the transmitted and received rf power at each frequency. It results in a transfer function

plot with an accuracy of ± 0.1 dB. The photodiode frequency response was provided by the

manufacturer and was subtracted from the result.

In addition to transmission magnitude (Fig. 6 (b)), the network analyzer also measures phase

response. The linear component of the phase due to device transit time is removed from the

measurement by entering a compensating electrical delay into the network analyzer. As plotted in

Figure 3, the phase response slowly degrades as the frequency increases, but no sharp deviations

are observed. Therefore, phase-related dispersion is minimal. If the phase response were poor, one could safely conclude that the device has a poor rise time.

2.3. FFT of the impulse response

The third measurement is performed in the time domain using a 15-ps step pulse.^ An illustration

of the experimental setup is shown in Fig. 4. The impulse response of the device is measured by

driving the rf electrode with a fast electrical step pulse whose amplitude is much less than Vji

(typicaJly 25%). This amplitude is chosen to keep the optical output of the device in its linear

region. We biased the device to about 75% of maximum on the transfer curve, and applied a negative-going step pulse having a fall-time of 15 ps and an amplitude of 5 volts (Fig. 5). The

resulting optical step was converted using a 20-GHz InGaAs photodiode, and captured on a 50-

122
GHz sampling scope. The derivative of the input step was then deconvolved from the derivative of

the measured output step to provide the true linear impulse response of the MZ. We obtained the

result in Fig. 6 (c) by transforming the impulse response to the frequency domain, and subtracting

off the photodiode frequency response.

3. SUMMARY

The swept-frequency system has two real advantages: 1) the measurement is insensitive to noise or

stray rf signals because it uses a lock-in amplifier, and 2) the measurement is independent of the

photodiode bandwidth because signal detection occurs at the 5-kHz modulating frequency. Due to the nature of this measurement, however, the phase response cannot be obtained. The network

analyzer technique is the only one that yields usable phase information and a direct measurement of

frequency response, rather than a calculated one. However, a disadvantage of this method is that

its results cannot be properly transformed to the time domain, due to the finite-width frequency

window used. The "FFT" method yields the actual time-domain impulse response of the device.

Thus, the response of the device to arbitrary inputs in the linear region of its operating curve can be

predicted. Due to the finite time window used to capture the impulse response, however, the FFT

may introduce some low frequency errors. Furthermore, differentiation of the input and output

waveforms contributes noise to the deconvolution process.

A summary of the advantages and disadvantages of each method is given in Table 1. All three

techniques produce similiar results (Fig. 6) and can be used to improve packaging or electrode

designs. For example. Fig. 7 shows how LLNL used such data to produce our current 1300 nm

devices with much smoother and well behaved frequency responses while decreasing Vtc to 10

volts.

4. REFERENCES

[1] R. C. Alfemess, "Waveguide electrooptic modulators," IEEE Trans. Microwave Theory

Tech., vol. MTT-30, No. 8, pp. 1121-1137, Aug. 1982.

[21 R. A. Becker, "Traveling-wave electro- optic modulator with maximum bandwidth-length

product", App. Phys. Lett. 45, No. 11, pp. 1168-1170, Dec. 1984.

[3] M. Izutsu, Y. Yamame, and T. Sueta, "Broad-band traveling-wave modulator using a LiNbOs optical waveguide," IEEE J. Quantum Electron., Vol. QE-13, pp. 287-290, Apr. 1977.

[4] M. Izutsu, T. Itoh, and T. Sueta, "10 GHz bandwidth traveling wave LiNbOs optical

waveguide modulator, " IEEE J. Quantum Electron., vol. QE-14, pp. 394-395, June 1978.

[5] S. K. Korotky and R. C. Alfemess, "Time- and Frequency-domain response of directionalcoupler traveling-wave optical modulators", J. Lightwave Tech., Vol. LT-1, No. 1, 244, pp. 244-

251, Mar. 1983.

[6] S. Uehara, "Calibration of Optical Modulator Frequency Response with Application to Signal Level Control", Appl. Opt. 17, No. 1, pp. 68-71, Jan. 1978.

[7] K. Kubota, J. Noda, and O. Mikami, "Travehng wave optical modulator using a directional

coupler LiNbOs waveguide", IEEE J. Quantum Electron., Vol. QE-16, No. 7, pp. 754-760, July

1980.

[8] S. K. Korotky, G. Eisentein, R. S. Tucker, J. J. Veselka, and G. Raybon, "Optical

intensity modulation to 40 GHz using a waveguide electro-optic switch", App. Phys. Lett. 50, No.

23, pp. 1631-1633, June 1987.

Work perfonned under the auspices of the U.S. Department of Energy by Lawrence Livermore National Laboratory under

Contract No. W-7405-Eng-48.

123
GPIB HP8341

Swept Freq. Synthesizer

1300-nm FP

laser diode

DC

bias

PM

fiber

isolator PM

fiber

RF

out MachZehnder SM

fiber

lock-in

mod. ref. amplifier

in in

i isig.

in

f low-bandwidth f photodiode

Fig. 1 Experimental setup for swept-frequency measurements.

HP8510C Network Analyzer

GPm

RFin

1300-nm FP

laser diode

DC

bias

PM

fiber

isolator PM

fiber

1=1 1=1 m m Q S —o o-*

RFout

MachZehnder SM •

fiber

HP83440C20GHZ

[ photodiode

Fig. 2 Experimental setup for network analyzer measurements.

0 5 10 15 20 25 GHz

Fig. 3 Phase measurement of a device using the network analyzer.

1300-nm FP

laser diode

GPIB

HP54I24T50GHZ

Sampling Scope

DC

bias

Pulser TDRstep oD

PM

fiber

isolator PM

fiber

15-ps step

MachZehnder SM

fiber

Fig. 4 Experimental setup for impulse response measurements.

-40-

0 100 200 300 400

ps

Fig. 5 Step pulse with 15-ps fail time.

Table 1. Summary of Methods

Method Advantages Disadvantages

Swept

Frequency

Independent of PD

Bandwidth No Phase Info. No Deconvolution

Network

Analyzer Measures Magnitude

& Phase No Deconvolution

15-ps Step Actual Impulse Response Noisy FFT

^5- 8 «"

T3a1

s

« -55-

-50-

-60-

GHz

Fig. 6 Frequency response comparison using 3 different methods:

(a) Swept frequency

(b) Network analyzer

(c) FFT of impulse response

5CD 2

Fig. 7 Swept frequency response of a recent LLNL device with an improved electrode design.

124
FIBER DIAMETER FROM THE BACKSCATTERED LIGHT PATTERN

John Baines\ Catherine FrehilP, Ken Raine^

1

- National Physical Laboratory

Queens Road

Teddington

2- School of Physical Sciences Dublin City University

Glasnevin

Dublin 9

Ireland

Middlesex TVVll OLW

UK

ABSTRACT

We have used the method proposed by van der Mark & Bosselaar [1] to determine the

diameter of single mode fibers from the backscattered light pattern. Measurements of the mean

diameter of three fibers using this method agreed to 0.02pm with measurements using a scanning corvfocal microscope with interferometrically measured stage movement. A fourth fibre

with large nondrcularity demonstrates the limitation of the method.

BACKGROUND

Marcuse & Presby [2] investigated light scattered by optical fibers with arbitrary refractive

index distributions. They found good agreement between measured light distributions and those

predicted by scattering theory. They noted, however, that the observed backscattered pattern

was extremely sensitive to changes in diameter and refractive index. The so-caUed inverse

scattering problem, that of determining the properties of the scatterer from the observed

scattering pattern, is complex. In the present case the backscattered pattern has periodic

solutions which depend on both the diameter and the refractive index of the fiber, van der Mark

& Bosselaar proposed observing the backscattered pattern at two wavelengths to remove the

ambiguity due to uncertainty in the refractive index.

All of the fibers measured here were single mode; van der Mark and Bosselaar have

shown that the effect of single mode cores on the backscatter pattern is insignificant. The values

for the fitted refractive index confirm earlier results [3] that the cladding index is raised during

the fiber drawing process by about 0.000 4. For the dispersion from 543nm to 633nm we have

used the known value for bulk silica. To remove the ambiguity caused by multiple solutions

the approach we have used therefore is to assume the refractive index of the sample is known

to within about 0.000 02 and the diameter is known to about 0.2pm.

Light from the laser was expanded then coUimated and passed through a wedged beam

splitter with low and high reflectance coatings on the front and rear surfaces respectively. The

laser sources were plane polarized and in the experiment the plane of polarization was rotated

using a half-wave plate. The fiber was perpendicular to the beam and backscattered light was

reflected by the rear surface of the beam splitter onto a linear CCD array with 2048 pixels. The

intensity profile across the backscatter pattern between about -5*^ and +15° was then digitized

and stored on the computer. We measured backscatter patterns from four telecommunications

METHOD

125
fibers with nominal diameters of 125]im at a wavelength of 633nm and two of these fibres at were also measured at 543nm.

We have calculated backscatter patterns using Bohren and Huffman's program [4]

(modified by van der Mark) for fibers in tiie diameter and refractive index ranges shown in the

table below and then searched by eye for the best fit between the observed and calculated

patterns. The patterns were calculated for light with the plane of polarization parallel (p) and

perpendicular (s) to the plane of incidence.

Laser

wavelength

Diameter

range (pm)

Refractive index

range

633nm 124.0 - 126.0 1.457 2 - 1.457 6

543nm 124.0 - 126.0 1.460 0 - 1.460 5

Increment 0.01 pm 0.000 1

We found that it is essential to pay particular attention to minimizing stray reflections

from the optical components and fittings. The wedged beam splitter, its coatings and the beam

dump are also important to make sure the recorded backscatter pattern was not distorted.

RESULTS

We have compared the results from the backscatter method with measurements made at NPL using a scanning confocal microscope (SCM) [5] and also, for three of the fibers, at NIST

using a contact micrometer [6]. The results at 633nm are shown in the table below together with

the estimated uncertainty in the result. Note that fiber 4 was not measured by NIST.

Fiber # Diameter

(pm)

Refractive

index

NIST

Micrometer

value (pm)

NPL

SCM

value (pm)

1 125.03 1.457 3 125.02 125.00

2 125.56 1.457 4 125.55 125.55

3 125.11 1.457 4 124.91 124.86

4 125.25 1.457 3 125.25

Estimated

Uncertainty

(95%)

0.03pm 0.000 1 0.03pm 0.05pm

126
Fibers 1 and 2 were also measured at 543nm and the results are also shown in the table

below.

Fiber # Diameter

(pm)

Refractive

index

1 125.03 1.460 5

2 125.54 1.460 5

The values for the measured diameter are the means of the results obtained for the two

polarization states. Since the fibers are elliptical rather than circular the results given are the

means of measurements with the fiber at two orientations 90° apart.

DISCUSSION

The tables show that, except for fiber 3, the three methods show excellent agreement (less

than 0.02pm). Fibers 1, 2 and 4 have low noncircularty (less than 0.3%) but that of fiber 3 is significant (about 0.8%). The SCM and micrometer results for this fiber differ by O.OSpm but the

backscatter method measures about 0.2pm too high. For a constant refractive index the

backscatter pattern almost repeats every 0.18pm at a wavelength of 633nm. Since the differences

in the pattern between one solution and the next are subtle they can be masked by instrumental

effects. It is therefore possible that the periodicity of the solutions can account for this result.

No fiber is exactly circular in cross section but deviations from circularity are nearly

always elliptical. For all cases we have observed, as the fiber is rotated perpendicular to the

coUimated beam the backscatter pattern changes but is always symmetric about 0°. Note that

the program assumes circular symmetry. However, the pattern at a given orientation seems to be the same as would be seen from a fiber with the same diameter as the projected diameter and

so measurements on elliptical fibers can be made. The search procedure is simplified

considerably if the diameter of the fiber is known within about 0.2pm. If a fibre has a

noncircularity greater than one period in diameter space then an ambiguity may arise.

Because the fitting procedure by eye can be tedious and the goodness of fit is difficult to quantify one of us (FrehOl [7]) has investigated using various techniques to automate the

process and quantify the fitness of the solution. Genetic algorithms have proved suitable for

efficiently solving complex optirruzation problems. They are highly parallel search procedures

which follow simple rules to mimic the processes of evolution by natural selection to converge

to an optimum solution. Preliminary results using noiseless test data show that, with a proper

choice of fitting parameters, genetic algorithms can successfully search refractive index and

diameter space to find the best fit to the backscatter pattern. Whether or not the method is practicable with real data remains to be seen.

CONCLUSIONS

We have implemented the proposal by van der Mark and Bosselaar to use the measured

backscatter pattern to infer the diameter of optical fibers. With careful experimental design the

127
results obtained agree with other reference methods to less than 0.02pm for fibers with low

noncircularity. The approach we have used requires that approximate values for the refractive

index and diameter are known but these values are easily obtainable. The method is perhaps

most suitable as an independent check of other methods used for calibration. The method has

the advantage of needing no calibration other than the wavelength of the laser provided the

refractive index is known. The four fibres we measured had an index within 0.000 05 of 1.457

35. Measurements at a second wavelength can be used to verify the diameter.

REFERENCES

[1] van der Mark, M. B., Bosselaar, L., "Fibre diameter calibration with lOnm accuracy

potential". Conference Digest, OFMeC / COST 217 Conference on Optical Fibre Measurements

held at York, England, September 1991, pp 147 - 150.

[2] Marcuse, D., Presby, H. M., "Light scattering from optical fibers with arbitrary refractive

index distributions", J. Opt. Soc. Am., vol 65, no 4, pp 367 - 375, 1975.

[3] Raine, K. W., Baines, J. G. N., King, R. J., "Comparison of refractive index measurements of

optical fibres by three methods", lEE Proc, vol 135, part J, no 3, pp 190 - 195, 1988.

[4] Bohren, C. K., Huffman, D. R., "Absorption and scattering of light by small particles", Wiley

Interscience, New York, Appendix C, 1983.

[5] Collyer, A., Raine, K. W., Baines, J. G. N., "Investigation into the use of a scanning confocal

microscope as a primary method for determining optical fibre diameter". Conference Digest,

OFMeC / COST 217 Conference on Optical Fibre Measurements held at York, England,

September 1991, pp 143 - 147.

[6] Young, M., Hale, P. D., Mechels, S. E., "Optical Fiber Geometry: Accurate Measurement of

Cladding Diameter", J. Res. Nat. Inst. Stds. Tech., vol 98, no 2, pp 203 - 216, 1993.

[7] Freehil, C, "Backscatter Analysis using Genetic Algorithms", Final Year Report, Dublin City

University, Ireland, 1994.

ACKNOWLEDGEMENTS

The authors would like to thank Martin van der Mark of Philips Research, Eindhoven for

suggesting this work and providing the software to calculate the backscatter patterns. We would

also like to thank Matt Young of NIST for permission to use the micrometer data.

128
FIBER GEOMETRY: RESULTS OF AN

INTERNATIONAL INTERLABORATORY MEASUREMENT COMPARISON

Timothy J. Drapela, Douglas L. Franzen, and Matt Young

United States Department of Commerce

National Institute of Standards and Technology

Mail Stop 814.02, 325 Broadway, Boulder, CO 80303-3328

Improvements in technology have led to a tightening of tolerances for optical fiber geometrical

parameters. This, in turn, has led to the need for improved measurement accuracy. Of the geometrical

parameters, cladding diameter accuracy has been technically the most difficult to improve. Accuracy on the order of ±0.1 |am is needed for the industry to comfortably meet specified tolerances of ±1.0 jam. In order to compare results with previous interlaboratory comparisons, we report average

measurement spread per fiber, for each measured parameter. We obtain this number, for a given parameter,

by first calculating the sample standard deviation for each fiber. We then calculate the arithmetic average of

these standard deviations. Although this is not a statistically valid overall measurement spread, it does give an indication of the relative agreement among participants. In 1 989, the Consultative Committee on Intemational Telegraph and Telephony (CCITT, now the

International Telecommunication Union - ITU) completed and reported on an intemational interlaboratory

comparison of fiber geometry measurements.' The average measurement spread per fiber for mean cladding

diameter was 0.38 |im and deemed to be unacceptably high. Each participant received a separate set of fiber specimens, so measurements were not made on the same fiber ends. Longitudinal nonuniformity could have

therefore influenced the results. The high measurement spread pointed to the need for all measurements to be made on the same cleaved fiber ends and also to the desirability of some type of calibration artifact.^ In North America, a second interlaboratory comparison was coordinated by the National Institute of

Standards and Technology (NIST) for the Telecommunications Industry Association (TIA) and reported on

in 1992.^ Participants serially measured the same cleaved fiber ends, which could be retracted into an aluminum housing for protection between measurements and during shipping. These housings use a brass

barrel that is moved to extend or retract the fiber and is marked to identify the angular orientation of the

fiber. The housed fiber specimens became the basis for the NIST calibration artifact or Standard Reference

Material (SRM), which became available to the industry in late 1992.'' The average measurement spread per fiber for mean cladding diameter measurements in the 1 992

comparison improved to 0.15 |am. There were also reductions in the measurement spreads of the other

measured parameters: cladding noncircularity and core/cladding concentricity error. Participants' cladding

diameter measurements were also compared to measurements made by the NIST contact micrometer, which

is the measurement method for characterization and certification of the SRMs. The contact micrometer is accurate within ±0.045 |im (3 standard deviations) for mean cladding diameter measurements on the housed

specimens.' The cladding diameter measurement spread observed in the 1992 study was largely due to systematic differences between participants, and fiirther improvement in this spread was anticipated with the use of calibration artifacts such as the NIST SRMs.

The present ITU intemational comparison was administered by NIST, with help fi-om regional

coordinators in Europe and Japan. Housed fiber specimens were again used. Because the housed fibers could become damaged with repeated shipping, and in order to complete this comparison in a reasonable

time, three sets of specimens were prepared for three different regions: North America (A), Europe (E), and

Pacific (P). Each set contained seven housed fiber specimens, including both very nearly circular and

noncircular (up to 1.5 %) fibers. All fibers were measured by the NIST contact micrometer, to allow for

inter-regional comparison. As of this writing, twenty-two participants have reported results; the number

129
should approach thirty by completion of the comparison. Of the current twenty-two, twenty used the gray

scale measurement method,* and two used image shearing.^ Of the gray scale participants, three used homemade test sets, while the others used commercial test sets from three different commercial vendors.

Table 1 shows the average measurement spreads per fiber, for each of the three measured parameters,

for the present as well as the previous comparisons. It also breaks down the present results into the three

regions. With the use of the housed fiber specimens, the spreads for cladding noncircularity and for core/cladding concentricity error have improved significantly for all three regions since the 1989 comparison.

In North America, where the NIST SRMs are in widespread use for calibration, the average spread per fiber for mean cladding diameter measurements has again improved, to 0.08 jam. The European and Pacific

spreads remain higher, both at >0.30 although both deserve quahfication. In Europe, many of the

participants are now calibrated to the National Physical Laboratory (NPL) in the United Kingdom, through a program similar to the NIST SRMs,^ involving either fiber or chrome-on-glass artifacts. If we include data

from only those participants, the European spread improves to 0.07 pm. In the Pacific, only two Japanese

participants are cahbrated to a national standards laboratory, by use of fibers certified by the Japanese Quality

Assurance Organization (JQA). Furthermore, a couple of Pacific participants (including one of the abovementioned calibrated Japanese participants) reported compatibility problems between the fiber housings and

their test sets, so they could not easily and confidently measure the housed fiber specimens. In a separate

comparison, on three different non-housed fiber specimens measured by the NIST contact micrometer and the

Pacific participants, a corresponding spread of roughly 0. 1 1 ^m was obtained. Also, while virtually all data

from North American participants have been received, there are several European and Pacific participants,

from whom we anticipate data, who have not yet received or measured the comparison fibers. Figure 1 shows the cladding diameter results for the present comparison, as a plot of offsets from

NIST contact micrometer values versus participant. Up to seven points are plotted per participant,

representing the seven fibers in each set of measurement specimens. The participants are grouped by region.

Filled-in circles show gray scale participants who are calibrated by means of a NIST SRM. Filled-in squares show gray scale participants who are calibrated to NPL. Filled-in diamonds show gray scale participants who

are calibrated to JQA. Open stars show gray scale participants who reported compatibility problems between

their test sets and the fiber housings. Other gray scale data are denoted by open squares. Open triangles

denote the two participants who used the image-shearing method. This graph clearly shows that those

participants who are calibrated to one of the national standards laboratories are usually in better agreement

than those who are not; in some cases, the improvement is nearly an order of magnitude. Agreement

generally appears to be just as good for noncircular fibers as for circular ones. One unexplained observation,

presently under study, is that most of the participants calibrated to one of the national standards laboratories have positive offsets from the NIST contact micrometer. In other words, they systematically measiare slightly

higher than the contact micrometer.

Two other meaningful quantities can be calculated from the statistics of these measurement offsets. For each participant, an average offset (average of the seven plotted offset values) can be calculated, as can

the standard deviation of the seven offset values about that average. The average offset indicates systematic

offset from the NIST contact micrometer, and when compared to the same quantity for other participants, it indicates the extent of systematic disagreement between them. The magnitude of this quantity can be

minimized by calibration. The second quantity, the participant offset standard deviation or offset spread, is a

reflection of, among other things, the participant's random uncertainty; this value would not be expected to improve with calibration. For most of the participants who are not calibrated to one of the national standards

laboratories, the absolute values of their average offsets, what we call their offset magnitudes, are greater than

their offset spreads. Such participants could be anticipated to benefit significantly from calibration. All

participants who are calibrated to one of the national standards laboratories have offset magnitudes of

<0.16 ^m, and all but three have <0.1 ^m.

130
Table 2 shows the average participant offset magnitude and average participant offset spread from

the present comparison: overall, broken down into the three regions, and for only those participants who are

calibrated to one of the national standards laboratories. The same numbers from the 1992 TIA North

American comparison are also shown. In North America the average offset magnitude has reduced from

0.114 nm in 1992, when there were no SRM calibrations, to 0.073 )jm in the present comparison in which

nearly all of the North American participants are calibrated by means of SRMs. The European average

offset magnitude is somewhat higher, apparently due to the few European participants who are not calibrated

to NPL. If we include only those who are calibrated to NPL, the number reduces to 0.071 ^m. The

average offset magnitude is also higher in the Pacific, where only a couple of participants so far are

calibrated to any national standards laboratory. Again, the higher average offset spread in the Pacific is likely due to those test sets that had difficulty measuring the housed fiber specimens.

In conclusion, this comparison shows better agreement among participants, for all three fiber geometry parameters, than did the 1989 comparison. Where there are substantial disagreements, they seem

to be systematic; this was not always the case in the 1989 comparison, where there was more random spread

in the data. Those participants whose test sets are calibrated to one of the national standards laboratories,

through calibration artifacts, generally show significantly better agreement.

ACKNOWLEDGMENTS

Thanks to: all participating laboratories in the United States, Canada, United Kingdom, France, Italy, Finland, Netherlands, Sweden, Japan, and Australia; to John Baines of NPL for coordinating and overseeing the European measurements and to Masaharu Ohashi of NTT for coordinating and overseeing the Japanese measurements;

to Bill Kane and Tom Hanson at Coming, Inc. for keeping us aware of ITU goals and deadlines and for reporting the

results; to Casey Shaar of Photon Kinetics for providing the prototype design of the fiber specimen housings; to Steve Mechels of NIST for his expert end-preparation on the measurement specimens.

REFERENCES

1. CCITT Study Group XV, "Results of a Round-Robin Study into Measurement of the Geometrical Properties of SingleMode Fibres," September, 1989.

2. John G. N. Baines, Andrew G. Hallam, Ken W. Raine, and Nick P. Turner, "Fiber Diameter Measurements and Their

Calibration," J. Lightwave Technol, vol. 8, no. 9, pp. 1259-1267, September, 1990.

3. Timothy J. Drapela, Douglas L. Franzen, and Matt Young, "Single-Mode Fiber Geometry and Chromatic Dispersion:

Results of Interlaboratory Comparisons," Technical Digest: Symposium on Optical Fiber Measurements, 1992, Nat. Inst. Stand. Technol. Spec. Publ. 839, ed. by G. W. Day and D. L. Franzen, pp. 187-190, September, 1992.

4. National Institute of Standards and Technology, Standard Reference Material Program, Bldg. 2, Rm. 204, Gaithersburg,

MD 20899; (301) 975-6776. Refer to SRM 2520, Optical Fiber Diameter Standard.

5. Matt Young, Paul D. Hale, and Steven E. Mechels, "Optical Fiber Geometry: Accurate Measurement of Cladding

Diameter," /. Res. Nat. Inst. Stand. Technol, vol. 98, no. 2, pp. 203-216, March-April, 1993.

6. Fiberoptics Test Procedure FOTP-176, "Method for Measuring Optical Fiber Cross-Sectional Geometry by Automated

Grey-Scale Analysis," Telecommunications Industry Association - Electronic Industries Association, 2001 Pennsylvania

Avenue, NW, Washington, DC 20006.

7. M. J. Downs and N. P. Turner, "Application of Microscopy to Dimensional Measurement in Microelectronics," Proc. Soc. Photo-Opt. Instrum. Engrs., vol. 368, Microscopy-Techniques and Capabilities, pp. 82-87, 1982.

8. J. Baines and K. Raine, "Review of Recent Developments in Fibre Geometry Measurements," Technical Digest:

Symposium on Optical Fiber Measurements, 1992, Nat. Inst. Stand. Technol. Spec. Publ. 839, ed. by G. W. Day and D. L. Franzen, pp. 45-50, September, 1992.

131
0.6

0.5

0.4

0.3

0.2

0.1

0.0

-0.1

-0.2

-0.3

-0.4

-0.5

-0.6

-0.7

-0.8

North American Region

••••

# Gray Scale: calibr. to NIST SRM

Gray Scale: callbr to NPL ^ Gray Scale: callbr to X!A

ir Gray Scale: test-setA^ousings Incompatibility Q Gray Scale: oltier A Image Shearing

1—I—

5tii

European Region Pacific Region

At

1 2 3

Figure

plotted,

1

\—\—I\—\—I

\—\—I\—I—I\ \ r

7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22

Participant number

.1 Offsets of participants' mean cladding diameter measurements from NIST contact micrometer. Up to seven points

for each participant, represent the seven fibers in the measurement samples.

Average Measurement Spreads per Fiber

Cladding

Diameter,

^Jm

Noncircularity,

%

Concentricity

Error, pm

'89 CCITT

Comparison 0.38 0.27 0.17

'92 TIA

(North American; Comparison

0.15 0.05 0.04

Present Comparison

0.24 0.09 0.08

"A" 0.08 0.05 0.06

"E" 0.30 0.11 0.12

"P" 0.33 0.12 0.07

Table 1 Average measurement spreads per fiber (arithmetic average of standard deviations for each of the

fibers), for all three measured parameters, for the current comparison, as well as past comparisons. Current results are also broken down by region: North America (A), Europe (E), and the Pacific (P).

Average Statistcs for Participants' Cladding

Diameter Offsets from NIST Contact Micrometer

Average

Offset Magnitude,

|jm

Average

Offset Spread,

|jm

'92 TIA

(North American) Comparison0.114 0.032

Present Comparison

0.171 0.052

"A" 0.073 0.027

"E" 0.188 0.048

"P" 0.299 0.101

those calibrated

to nat standards

labs (all regions) 0.069 0.030

Table 2 Average cladding diameter offset statistics per

participant for current comparison (overall, regional, and only

those calibrated to national standards laboratories) and '92 North American comparison. Average offset magnitude is the average of the absolute values ofparticipants' average offsets from NIST

contact micrometer. Average offset spread is the average of the

Participants' standard deviations about their average offsets.

132
Contact and non-contact interferomefry for optical fiber diameter

measurement and calibration

Claude Brehm, Olivier Mercereau, Genevieve Lipinski *, George Pierre Vailleau *.

Alcatel Alsthom Recherche UFP, route de Nozay 91460 Marcoussis, France

• LNE Laboratoire National d'Essais, 1 rue Gaston Boissier 75015 Paris France

Introduction

• Fiber outside diameter measurements must be made on-line during fiber drawing in order

to be as close as possible from the nominal value of 125 )im on the whole fiber length.

On-line measurement techniques which are typically used in fiber manufacturing units are

laser scanning fiber shadow equipments. Tnese are very reliable and quite convenient

for a close-loop control on capstan speed during fiber drawing. However they show two

main drawbacks :

- 1) they are quite slow and cannot detect localized fiber diameter variations at high

drawing speed.

- 2) calibration is made with metallic wires which are known with a limited accuracy

(125 ±0.2 jim) and give a systematic discrepancy with fiber end cut measurement.

Therefore a dynamic-in^erferometry device for hign speed measurement of fiber diameter

fluctuation has been promoted and a new non-contact interferometry static measurement

device has been build and tested ^

.

This second equipment can be used to characterise standards for the calibration of

dynamic interferometer and other diameter measurement devices.

The purpose of this presentation is to describe more extensive tests which have been

made both in Alcatel Alsthom Recherche and in Laboratoire National d'Essais (LNE)

where a comparison with contact interferometry measurements has also been made . Contact interferometry for fiber diameter measurements has already been described and

compared by various laboratories such as NIST , NPL ^ , and Corning ^.

Contact Interferometry for measurement of outside diameter of optical fiber

• Principle of measurement

Our investigations concerning the choice of a principle of measurement for the

determination of outside diameter of optical silica fiber lead us to select a measuring

method allowing a direct traceability to the definition of the meter. The method is based on an interferometric technique using the fringe fraction measurement.

A Michelson interferometer lighted by two stabilised He-Ne laser sources : a green one

[\= 543 nm) and a red one (A.= 633 nm), which was used for the calibration of gauge

blocks, has been modified to measure the diameter of small cylinder. The fiber is inserted

between two gauge blocks of 10 mm of nominal length. The lower gauge is wrung on one of the metallic optical flat mirror of the Michelson interferometer. The upper gauge

block is supported by a small sphere and by the fiber (along a generating line). The

parallelism of the upper gauge with the optical mirror is ensured by vertical motion of the

supportina sphere induced by piezo actuators. The interferential image is composed by two fringe patterns, one is produced by the

reflection on the surface of the upper gauge and the second is produced by the reflection on the mirror. A CCD camera digitalize this image and an image processor gives the

fringe fraction. The length of the two gauges has been measured previously with the same

method. The assembly composed by the gauges and the fiber is then measured

[h = (hl+h2)/2] and the fiber diameter is deduced (figure 1 side and top views).

133
• First results and estimation of measuring uncertainty.

This method allows to measure a mean outside diameter on a fiber length of 9 mm.

It is also possible to repeat measurement, after rotation of the fiber around its axis, in order to estimate fiber non-circularity. This contact method is independent of optical

parameters of the fiber. The measuring uncertainty has been estimated by using the method described in the guide

to the expression of uncertainty in measurement, recently published by ISO

The different sources of uncertainty have been consiclered as independent, so that the

combined uncertainty is : Uc =47 nm ( see table 1) and the expanded uncertainty, with a

covering factor of k = 2 is : U = ± 94 nm. The analysis of different sources of uncertainty shows that the main component is repeatability. It is estimated by the standard deviation of a set of successive

measurements after rotation of the fiber on its axis. It is mainly limited by two factors : the

repeatability of the measuring process itself and the geometrical defects like fiber noncircularity. Presently the main limiting factor for the uncertainty is the geometrical quality

of the fiber which is now being improved.

Non-contact interferometry : scattering pattern from a side-illuminated optical fiber. • As shown in figure 2, two rays are involved, Re reflected at the fiber surface and T

refracted into the fiber. These two rays do not have the same incident angle i, therefore

interference appears. If D is the fiber diameter, 9 the scattering angle and n the refractive

index of the fiber, the optical path difference, d, is : 3 = D. ( sin 9/2+ (n2+l-2n cos 9/2)1/2).

Between two scattering angles the number of interference fringes is only dependent on

the fiber refractive index and on the fiber diameter. This gives a direct measurement of the

fiber diameter if the fiber index is known.

For single-mode fibers the core diameter is less than one tenth of the fiber outside

diameter and therefore the core-cladding index difference has no effect in the scattering

angle range used for measurement, i.e. 9 between 10 and 80° . • A schematic layout of the measuring equipment is shown in figure 3. The beam of

polarised He-Ne laser is focused on the fiber which is placed at the center of a goniometer. Polarisation and fiber axis have the same direction. The scattering pattern is focused on the photo detector which is set on a motor-driven rotation stage. A crossed

polarizor is placed in front of the photo detector to avoid saturation in the direction of

incident ray. Measurement is made on a fiber height of 100 |j.m. Power data is collected

every 0.01° for the determination of fringe pattern limits and every 0.03° for fringe

counting. Averaging is done on 2000 sample points at 20 kHz. Measurement accuracy

then is better than 0.1 |.im.

Comparative measurements

• Measurements have been made on a typical single-mode fiber, repeatability is less than

10 nm (Istandard deviation). Five successive sets of 100 measurements on a 10 mm

fiber's height have been made, each one after a rotation of 72° of the fiber on its axis

(mean value : 125.13 jim). The data, minimum and maximum values for each angle, is gathered on figure 4; it shows a fiber non-circularity close to 0.2 )im. These data are

consistent with diameter measurements made by fiber end cut with a York S 20

grayscale optical microscope system (mean value : 125.14 |a,m). They are also very

consistent with measurements made at the French National Standard Laboratory (L.N.E.)

with a two wavelength Michelson interferometer (mean value : 125.14 |j.m). Here also,

measurement repeatability is low, due to the fiber non-circularity

• A fiber with a lower non-circularity has been taken for a second set of comparative

measurements Data are shown on figure 5. Fiber non-circularity has been check with the

two interferometric methods by rotating the fiber on its axis in 5 different directions, it is then close to 0.05 [im. Mean values are 1 24.90 (xm with contact interferometry (LNE)

and 124.81 |i,m with non-contact interferometry.

134
ConcI usion

• Data about outside diameter, obtained out of pieces of the same fiber close from one to anotfjer, are very similar and in the range of uncertainty. Slight geometrical imperfection

mainly nori-circularity, remaining in the fiber tested, is still limiting measurement uncertainty.

Better quality fiber could be useful for an even more accurate comparison of these two

measuring techniques.

• For measurement of outside diameter of optical fiber, interferometry v/ith two

mechanical contacts is the only method to reach a high accuracy without any need of

data about the effective refractive index of the fiber under test. Presently the main limiting

factor for the uncertainty is the geometrical quality of the fiber, such as surface defects

and non-circularity. Length of the mechanical contact which is presently of 9 mm could

also be reduced to improve the measurement accuracy.

• The non-contact static interferometry equipment has the ability to become, in each fiber

producing unit, a secondary standard maker for a more accurate calibration of various

fiber diameter measurement devices. Silica fiber of appropriate dimension can be used

for the calibration of various laser scanning shadow measuring heads instead of metallic

calibration wires.

Acknowledgements.

The authors would like to thank H. Gagnaire (Universite de St Etienne), G. Creel,

P. Darbon, P. Dupont (Alcatel Fibres Optiques), B. Minvielle (Alcatel Alsthom 'Recherche).'

References

[1] C. Brehm, P. Darbon, B.Minvielle, H. Gag noire "New technique for optical fiber

outside diameter measurement and calibration" IWCS 1993 pp 817-821

[2] M. Young, P. Hale, S.Mechels "Accurate measurements of fiber cladding diameter"

Symposium on optical fiber measurements 1992 pp 55-58.

[3] J. Baines, K. Raine " Review of recent developments in fiber geometry measurements"

Symposium on optical fiber measurements 1992 pp 45-50.

[4] K.A. Emig "A comparison of interferometric techniques for fiber cladding diameter

measurements" Fiber and integrated optics 1990 vol.9 pp 375- 379.

Sources of uncertainty Value of standard

uncertainty u(xi) A 1 ; Repetabilty 45 nm

B 1 : Air wavelength) 2 nm

B 2 : Wringing effect 1 .7 nm

B 3 : Fringe fraction 1 .7 nm

B 4 : Pfiase correction 3,3 nm

B 5 : Optical defect of tfie interferometer 3.3 nm

B 6 : Correction of the expansion 4.6 nm

B 7 : Expansion coefficient 2.7 nm

B 8 : Fiber compression 1 ,4 nm

B9 : Lengtfi of the upper gouge 6 nm BIO : Length of the lower gouge 6 nm

B 1 1 : Local flatness of gouge blocks 8 nm

Table 1

"IT

hi JZ

Figure 1 : Fiber O.D. contact interferometric

teclinique used at LNE.

135
Wavefront Re Ray

Figure 2

Fiber cross-section : optical path difference

between two rays Re and T scattered

in the same direction

Goniometer

Beam

expander Direct beam

|He-Ne Laser~t=

Top view

Optical

fiber

etector

Interference

lines

|He- Ne Laser

Side view

Rgure 3 : Rber O. D. Static measuring device based interference analysis

125.4

E 125,3

' 125,2

~1 ^ ^ ^ ^ ^ ~

100 measurements on 10 mm

non-contact Intorlerometry maxi- min.

Contact

interterometry measurements

LNE

Fiber end cut

- S20

0 72 144 216 288 360 Angle {'C)

Figure 4 : O.D. measurements on the same fiber

• by end cut ( Yorl< S20 ) • by contact interferometry ( LNE)

• by non contact interferometry with 5 rotation angles of the fiber on its axis

Fiber O.D.dim

125,0

124,9

124,8

124,7

124,6

Contact^

Interfere metry measurement

inter lerometry

(mm)

Figure 5 : O.D. measurements on the same fiber

• by non contact interferometry

( each 0.2 mm on a 10 mm fiber length, i. e. 50 points and with 9 rotation angles of the fiber on its axis, i. e. a total of 450 points ) mean value : 124.81 Mm - minimum ; 124.77 )im - maximum 124.86

• by contact interferometry ( LNE) on a 9 mm fiber length mean value : 1 24.90 nm - non-circularity - 0.05 ^m

136
Light Scattering Fiber/Coating Eccentricity Measurements for Thin Coatings

D. H. Smithgall

AT&T Engineering Research Center

Box 900

Princeton, New Jersey 08542-0900

I. Introduction

The concentricity of 125 /im fiber in a typical 250 /xm polymer coating is determined by

measuring the symmetry of a forward scattered light pattern, which appears as a well-defined

bright bar, about an illuminating laser beam [1]. The distinct edge to this bright pattern

results from the concentration of light near the fiber/coating interface. As the ratio of coating

diameter to fiber clad diameter decreases, the pattern width increases and the contrast decreases until for coating diameters below 185 /xm the bright bar pattern is no longer

distinguishable.

The lack of concentrated light energy in the scattering patterns produced by thin coatings

results in the emergence of an interference pattern produced by Ught refracted through the

coated glass and light reflected from the surface of the coating. This pattern corresponds to that used to measure clad fiber diameter [2]. Since all of the refracted light passes through

the fiber, the interference pattern wiU be influenced by fiber eccentricity, as weU as refractive

index and coating diameter. Thus, a measure of fiber eccentricity is available by comparing

the fringe patterns on either side of the optical axis.

n. Ray Trace Model

Referring to Figure 1, a fiber which is eccentric within the polymer coating can be modeled as two concentric circles with the center of the fiber ofliset from the center of the coating a

distance, d, at an angle, <j>, with respect to the optical axis. The diameter of the fiber is 2a and

the diameter of the coating is 2b. Assume that this structure is uniformly illuminated by a

collimated coherent laser beam. The intensity of the forward scattered light pattern at any

given angle, 6, results from the interference of light refracted through the structure and light

reflected from the surface of the coating. When the outer layer of this structure is much

larger than the center circle, some light passes only through this outer layer, and the diameter

of this outer layer can be determined by a simple expression [3]. In addition, the eccentricity

of the fiber can be determined from light which passes through both the fiber and coating,

projecting a bright bar into the forward scattered pattern [1]. When the coating is thin, eg. a/b > 1/n, aU light passes through the fiber, and only the forward scattered interference

pattern exists.

Referring to Figure 1, the eccentricity of the fiber within the coating can be described in terms

of an offset amplitude, d, and orientation with respect to the optical axis, </». The path of a hght

ray refracted through this structure is determined from the equations:

X = sin(Qo) = n sin(Q;i) 0 < ;c < 1 (1)

71 = sin(a2) = (b sin(ai) - d s'm(QQ -ai + 4>)) / a

137

(2)
72 = « 71 /ng (3)

sin(a5) = n sin(Q;4) = {a sin(a2) + d s\n{oiQ - + 2 ( ^2 " <^3 ) + <^ )) / ^ (4)

WTien the fiber is slightly eccentric, eg. when d/b < < 1, 7j, '-{^ and (4) are all less than 1, and

the refracted ray scattering angle is

^ = oo - ai + - + 2 (a2 - 03) . (5)

The reflected ray which scatters at the same angle, 9, is described by

9 = 7c-2fio (6)

The optical path lengths of the two rays between two planes normal to the entrance and exit paths are determined for the refracted ray by:

qQ = Z)cos(ai) - a cos(a2) + d cos{aQ -cti + <f>) (7)

qi = b cos(a4) - a cos(a2) - d cos(q^) - + 2(0:2 - ^3) + <f>) (8)

Pi = n {qo + qj) + rigla cos(a3) , (9)

and for the reflected ray by:

P2 = b (cos(q^) - cos(/9o)) + b (cos(a5) - cos(;9o)) (10)

If two angles are selected in the forward light scattering region, 9^ and 9^, then the number of

interference fringes, N, between these angles is

A^ = (A,-A,)/A , (11)

where A is the wavelength of the incident radiation, and each A is the optical path length

difference, pj - p2' ^ch angle.

When two identical angle ranges are selected, symmetric about the optical axis of the incident

laser beam, the difference in the number of fringes - N" depends upon the amplitude and

orientation of the fiber/coating offeet, since the clad and coating diameters and material

refractive indices are identical for both angle ranges.

As an example of the model-predicted results, consider a 125 /im fiber, coated with a 170 /im

diameter polymer coating. The refractive indices of the fiber and coating are 1.4571 and 1.54,

respectively, at the HeNe laser wavelength of 0.6328 /xm. Figure 2 illustrates the fringe pattern

differences in the angle range 9^ = 38° to = 52°.

In order to test this method of measurement, a fiber sample was mounted in a test apparatus,

which is described in Section III. Fringe measurements in the range 38° - 52° on either side of

the optical axis were made for 9 angular orientations of the fiber. Computing the maximum

fringe difference from the data, and comparing to the modeled results in Figure 2 predicted a

5 /xm eccentricity. The fiber sample was also measured by immersing the sample in index

matching fluid and using a microscope filar eyepiece to measure the maximum side wall

138
variation. These readings differed 6 and 5 nm from the nominal 22.5 /im concentric wall

thickness, which is comparable to the light scattering measurements, within the microscope

measurement accuracy.

HI. Implementation

Figure 3 illustrates the arrangement of components used to measure the differential fringe

patterns. The coated fiber is orthogonally illuminated by a HeNe laser beam, such that

scattering patterns of substantially equal intensity are generated on each side of the optical

axis. The detectors are arranged symmetrically about the optical axis.

Each detector section comprises a lens which images the scattering pattern onto a line scan

array, located in the back focal plane of the lens. When properly aligned, the detected pattern

is stationary for axial and lateral movement of the coated fiber. The focal length and f/# of

the lens determine the image size, the distance from the fiber to the lens and the window

within which the fiber may be located without affecting measurement accuracy.

The output from each diode array is a sample-hold boxcar signal, scanned from the array at a

rate of 1000 images per second by the detector electronics circuit. The two fringe patterns are

separately processed in circuits which measure the total number of fringes, or equivalently the

total phase in the fringe patterns in the fixed angle range defined by the line scan detector [3].

V. Summary

A technique has been developed to measure fiber/coating eccentricity in thin coated fibers

using forward light scattering. It has been used for in situ process control, with average

eccentricities less than 3.5 /zm over multikilometer lengths of fiber.

Acknowledgement

To C. M. Schroeder and D. F. Marcille for their assistance in evaluating this measurement

technique.

References

1. D. H. Smithgall, R. E. Frazee, Jr., "Measurement and Control of Fiber Coating

Concentricity", BSTJ, Vol. 60, No. 9, November, 1981, pp 2065 - 2080.

2. L. S. Watkins, "Scattering from Side-IUuminated Clad Glass Fibers for Determination of

Fiber Parameters", JOSA, Vol. 64, No. 6, June, 1974, pp 767 - 772.

3. D. H. Smithgall, R. E. Frazee, Jr., "Improved Real-time Optical Fiber Diameter

Measurement System", Technical Digest of CLEOS '80, February 26, 1980, San Diego,

CA, pp. 10-11.

139
FIGURE 1 RAY TRACE THROUGH ECCENTRK:

COATED FIBER CROSS-SECTION

0.5-

OffMt Ampiitud* (urn)

Fring*

Otffcrcncc

-0.5-

•1.0-

OftMt

Orisntation

FK3URE2

FRINGE DIFFERENCE AS FUNCTION OF FIBER OFFSET AMPLITUDE AND ORIENTATION

Coaling Indtx 1.54

HeNa

Fringe Counting

Circuit

Imaging Lens I Ainiy [f^ <\r Circuit Line Scan Army

Fringe Counting

CiroHt

PC

FIGURE 3 MEASUREMENT OF FORWARD LIGHT SCATTERED PATTERNS

FROM THIN-COATED FIBER

140
Transverse Optical Scattering Measurements of the Core Radius

of High-A Fibers with 1-nm Resolution

Peter N. Saeta

AT&T Bell Laboratories, 600 Mountain Ave., Murray Hill, NJ 07974 and NIST/JILA,

Campus Box 440, University of Colorado, Boulder, CO 80309-0440

Benjamin I. Greene, David J. DiGiovanni, and Anne E. Miller

AT&T Bell Laboratories, 600 Mountain Ave., Murray Hill, NJ 07974

An optical technique for measuring the core radius of high-A optical fibers is proposed in

which a laser beam is scattered off the core of a fiber immersed in index-matching fluid. Using a

fiber segment that was pulled to form a near-field optical tip, the technique is used to measure

variations in the core radius of step-index fibers down to the scale of 1 nm.

Standard techniques for measuring the index profile of optical fibers rely on scanning a focused spot of visible light across the cleaved fiber face. Their resolution is therefore limited

approximately to the wavelength of visible light, or about 0.5 /im. For step-index fibers with

large relative index difference. A, and core radii of only a few micrometers, this limited

resolution results in a significant distortion of the measured profile. Much greater resolution

could theoretically be obtained with scattering or interferometric techniques, ideally to a small

fraction of the wavelength. Measurements at the nanometer scale could provide useful

information on draw-induced changes in the profile and could conceivably test theories that

attribute the abnormally high loss of high-A fibers ^-^ to fluctuations in the core radius.'^-'^ This paper describes a high-resolution scattering technique suitable for high-A step-index

fibers. The method consists of scattering a focused laser beam off the core of the fiber and

analyzing the fringes of the scattering pattern away from the forward direction. It is similar to a conventional light scattering technique^'^ used to provide real-time feedback to control the fiber

diameter during the draw, except that the fiber is immersed in index-matching fluid. This

effectively removes the cladding, vastly simplifies the scattering pattern, and permits much more

direct information about the core to be obtained. Analysis of the resulting scattering pattern

permits nanometer-scale fluctuations of the core radius to be observed.

Light scattering from fibers has been used for many years to monitor the fiber diameter

during the draw and provide feedback to minimize diameter tluctations.^-^ The theory for

scattering plane electromagnetic waves off infinite homogeneous dielectric cylinders and

cylindrical annuli has been discussed by many authors.^-^-^ Briefly, we assume the index profile

to be a piecewise-constant function of the radius, r, and equal to the cladding index, n^^^^^, for r >

b. Then the scattered electric field, E{<p), can be expressed as a cosine series in the azimuthal

angle, 0, having complex coefficients determined by boundary conditions at the interfaces

between successive annuli. To obtain spatial resolution along the fiber axis and to maximize the

signal with respect to scatter from the cell and the fluid, a lens is used to focus the laser beam at the fiber core, taking care that the beam waist, w, satisfies w » b. The power detected in an

angle A0 in the far field is then given by

insofar as the deviations of the incident wave from plane can be ignored.

While the sum in Eq. (1) is infinite in the order parameter v, the coefficients Cy fall off

rapidly for v>2m^^^^^b^, =N. For the case of scattering from standard fibers in air, /V= 620,

Abstract

Introduction

Theory

(1)

141
which produces a pattern of very high

angular frequency dominated by the aircladding interface. The portion of this

pattern near the backscattered direction

has recently been used to measure

cladding diameter to -10 nm.'^ In the

present case of a nominally step-index

fiber with b ~ 1.2 /im, is approximately

15, so that the fringe spacing is roughly

180°/15= 12°.

Figure 1 shows the scattered power

calculated using Eq. (1) for a step-index

fiber using typical experimental

conditions. The dashed line shows how the

pattern changes for a 1% increase in b.

Clearly if stray light can be suppressed and

the pattern around (f) = 90° can be

measured with even modest sensitivity,

resolution better than 12 nm can be

achieved in these fibers.

(P (Degree)

Fig. 1 Calculated scattering pattern for a step-index fiber (A = 0.023, b=l.2 fjm) using 1 W of O.^M.^i-^mi light focused to 2w = 40 /Jin. A slit width of 0.25° is assumed. Tiie dashed hne

shows the modified pattern for a 1 % increase in core radius b.

Experiment

The experiments were performed on optical fibers with their polymer jackets removed. The

setup was a slight modification of one previously described in detail. ^^-^^ The fibers were

suspended in a 2.3-cm-diameter silica cell filled with index-matching fluid. Either the OAS^-jim

or the 0.5145-/im beam of an argon-ion laser was focused by a 13.5-cm focal-length spherical

lens into the center of the scattering cell. The beam was vertically polarized (parallel to the fiber

axis); typical beam power was 1 W.

Two detection schemes were employed. In the first, the beam was chopped to permit lock-in

detection and the scattered light was detected with an unamplitled silicon photodiode mounted

on an optical rail atop a computer-controlled rotation stage at a distance of -20 cm from the

fiber. The acceptance angle of the detector was determined by an adjustable-width vertical slit mounted 10 cm from the fiber on the rotating rail; typical acceptance angles were 0.1-0.3°. In the

second, a cylindrical lens was placed near the cell and the collected light was imaged on a photodiode aiTay.

Results and Discussion

The fiber used for most experiments

(Fiber A) had a silica cladding layer and a germanosilicate core (18% Ge02),

yielding a nominal An of 0.030 (A = 2.1%)

and core radius of -1.2 fjm. When

immersed in index-matching iluid and - illuminated with either argon line, the |

scattering pattern was easily visible by eye ^ through angles exceeding 90° from the

forward direction and appeared symmetric

about the forward direction. When the

index-matching fluid and the cladding had

slightly different indices of refraction, the

pattern was modulated by shallow highfrequency tringes having a line spacing of

^ ^^^^^^^^ ^^^^^^^.^^ p.^^^. ^ a tl action Ot a degiee. step-index high-A fiber. Dashed curve is best fit for step-index

The scattering pattern measured usmg p,.of-,,e. ^^ij^ ^urve is the best fit, giving a core radius of 1.22 the single photodiode is shown in Fig. 2 ^m.

142
for the range 35° < 0<9O°. On the vertical

scale of diode voltage, the background

level detected when the fiber was removed

from the cell was between lO-"^ and lO^ V

over this range. The data agree

qualitatively with the calculated pattern of

Fig. 1, having roughly the same fringe

spacing, although the successive maxima

fall off more rapidly than in the calculation

for the idealized step-index core profile

(dashed curve). Qualitatively, the rate of

fall-off of the successive maxima is a measure of the broadening of the index

step. An attempt was made to fit the data

using a core index profile function with a smoothed transition from /Tcore "clad0.03

0.02

J I I L

r inm)

Fig. 3 Index profile as measured using the refracted nearfield method with a York model S14 fiber profiler (heavy line) and as obtained from the fit shown in Fig. 3 (thin line). Several forms were tried, and the solid curve in Fig. 2 shows the best fit obtained.

The fit follows both the spacing of the

minima and the decrease in the maxima with 0 fairly well; the discrepancy at the minima results

primarily from the finite slit width.

The index profile corresponding to the fit of Fig. 2 is compared in Fig. 3 to measurements

made using the refracted near-field method with a York model S14 fiber profiler. The profile

obtained from the fit has an approximate transition width of 0.1 /im, whereas the measured one is roughly eight times broader. The broadening is primarily due to the ~0.5-^m spot size of the

laser used to make the measurements, which necessarily smooths abrupt index changes. The

scattering pattern calculated from the measured profile shown in Fig. 3 falls off much more

rapidly with <p than the data and, unlike the data, shows minima of quite variable depth. It is likely, therefore, that the fit gives a much more accurate picture of the transition region from core

to cladding.

An array detector was used to investigate the cores of two fibers from the same preform

drawn at the same speed but at different tensions and temperatures. A cylindrical lens was used

to direct and compress the radially scattered light onto the linear diode array. The data in Fig. 4 were not corrected for the resulting nonlinear relation between pixel and angle. The scattering

pattern of the fiber drawn at lower tension

(higher temperature) decays more rapidly

with angle, indicating that the transition

region between core and cladding is broader than in the fiber drawn at higher

tension (lower temperature).

Simulations showed that the pattern is much more sensitive to the core radius, b,

than to A. They also displayed the

common feature that for small changes in b, 5b/boc-5(j)^J(l)^^^, where 0„„„ is the

position of a local minimum in the

scattering pattern. This suggests that even

if a satisfactory fit cannot be made to the

data, useful information about changes in the core radius can be obtained by

accurately measuring the position of one

or more minima in the pattern as a function of distance z along the fiber.

1000

Pixel

Fig. 4 Scattering patterns from two fibers differing only in draw conditions observed with a photodiode airay showing

that a higher draw temperature produces a more broadened

transition from core to cladding index.

143
0.01

0.00

-0.01

-0.02

-0.03

-0.04

-0.05

Ih- 0 • • e, g

I I I I

I

I I I I

I

I I I I

I

I I I I

I

c Minimum near 44° o Minimum near 53°

• Micrograph

s

I .... I

' ... I. ... I I .... I 50 100 150 200

2 (urn)

250 300 350

Fig. 5 Plot of the relative change in core radius along the length of the fiber in the neck-down region. The filled diamonds were taken from a digitized micrograph; the open

symbols were obtained from the relative shift in the position of

minima near 44° and 53°.

To investigate the sensitivity of this

information, a segment of Fiber A was

laser-heated under tension, causing it to neck down to a point and break in the

heated region. Scattering measurements

were then made of this "tip" starting in the

region of the fiber that had not been

heated. The relative change in radius, 5b/ b,

was calculated based on the minima near 0 = 44° and 53°, and the results are shown as

a function of z in Fig. 5. These are compared to a profile deduced from a micrograph taken of the neck-down region

of the fiber, based on the assumption that

the relative change of the core is equal to the relative change of the outside fiber

diameter. The agreement between the two

curves calculated from the shift in minima

is excellent—better than a part in 1000 in

the unmodified region (z < 100 //m), and

quite good throughout. Since b ^ 1.2 jum,

this implies that the technique can resolve fluctuations in core radius smaller than 1 nm.

It should be noted that the analysis assumes perfect circular geometry. Deviations from

roundness certainly exist at some level in fibers, and they will perturb the scattering pattern.

Accounting for them is a challenging theoretical problem. Nonetheless, Fig. 5 suggests that this

technique provides useful information on the nanometer scale.

Conclusions

The angular scattering pattern produced from a high-A, nominally step-index fiber immersed

in index-matching fluid has been recorded and shown to provide a sensitive measure of the core

radius. While attempts to deduce the exact index profile from the scattering pattern were

unsuccessful, small variations in core radius produce proportional shifts in the angular position

of shaip minima in the scattering pattern. These shifts were used to profile the neck-down region

of a fiber tip with resolution belter than 1 nm in the core radius.

Acknowledgments

We gratefully acknowledge R. E. Betzig, L. R. Copeland, D. M. Krol, C. R. Kurkjian, S. L. McCall, and W. A. Reed for help and useful discussions.

References

B. .1. Ainslie, K. J. Beales, C. R. Day, and .1. D. Rush, EEE .1. Quantum Electron. QE-17, 854 (1981).

S. T. Davey, D. L. WiUiams, D. M. Spirit, and B. .1. Ainslie, SPIE 1171,1 81 (1989). D. Marcuse and R. M. Derosier, Bell Sys. Tech. J. 48, 3217 (1969), D. Marcuse, Appl. Opt. 23, 1082 (1984). D. Marcuse and H. M. Presby, J. Opt. Soc. Am. 65, 367 (1975).

11. M. Presby, .1. Opt. Soc. Am. 64, 280 (1974). H. M. Presby and D. Majcu.se, Appl. Opt. 13, 2882 (1974).

J. A. Stratton, Electromagnetic Theory (McGraw-Hill, New York, 1941). H. C. Van de Hulst, Light Scattering by Small Particles (Wiley, New York

P. N. Saeta, to be published in Appl. Opt. (1994).

M. B. van der Maik and L. Bosselaai', J. Lightwave Technol. 12, 1 (1994). A. E. Neeves and W. A. Reed, Appl. Opl. 31, 2072 (1992).

2

34

56

7

89

10

1

12

1957).

144
Measurement of Fiber Core Position for Fiber Array

K. Ozawo , .V. Ogawa . H. Ishido and Y. Hatton

Sumitomo Electric Industries. Ltd.

1. Ta> a-cho. Sakae-ku. Yokohama. 244. Japan

1. ABSTRACT

A new measuring method of the offset of fiber core position in a fiber array was investigated, hi this

metliod, the offset of the fiber core position in the master array is measured in ad\ ance. and the sample

arra> is set face to face with the master arra>-. and measured the relative offset of its fiber core position.

Using this method, the core position can be measured in a short time and b>' a simple equipment

compared with the conventional wa> [l]. The repeatabilitv' of the measurement is under 0.05|ajii in the

case of single mode fiber, so we can apph' this method to the inspection of fiber arrays.

2. INTRODUCTION

The Planar Lightwave Circuits(PLC) are coming into practical use as passive optical parts such as

couplers, splitters, filters. Fiber arra> s are necessan,- for the connection between these optical parts and

fibers and the>' will pla\ a more important role in the future. One of the most important specification of the fiber arra> is the offset of the fiber core position at its

endface. That is because the offset has a direct influence on the coupling loss with another optical de\ ices. The offset of the fiber core position means the de\ lation from the proper arrangement of fiber cores in

the endface of a fiber arra> . Fig. 1 shows the example of this offset in the case of a 4-core fiber array. G is the bars center of the centers of the four cores C1.C2.C3. and C4. and is the straight line obtained

b\ the least squares method from these four centers. A new coordinate s> stem is considered, in which G

IS the origin. 1-^ is X axis, and 1,^-. the perpendicular of l^^ passing G. is Y axis. In this coordinates system

the proper positions of the core D1.D2.D3. and D4 are arranged on \^ with certain distance a. The offset

distance of the fiber core is defined as the distance d between this proper position and the actual center of

the core. And the direction of the offset is defined as A6 in dns coordinate system.

There is conventional method to measure the offset of fiber cores for a multi-fiber connector[l]. A

CCD camera, precision XYZ stages, and laser interferometers are used in this method. The sample

connector is fixed on the XY stages and the CCD camera on the Z stage gets the core image on the

endface of the cormector b>' launching white light from the opposite side of the fiber. The core position

can be obtained by the image processing of the CCD camera and the laser interferometers. Needless to

say. this method can be applied to fiber arrays, and the repeatability of measurement is within 0.04^un.

But the cost of the equipment is high and time for measurement is rather long. So we investigated a

simpler measuring method of the fiber core position.

3. PRINCIPLE

In this method, the sample array is set face to face with the master array, the fiber core offset of

which ha\ e been already measured. The master array and the sample array are set face to face and the

clearance between the both endfaces is kept in a certain distance d around 30,um. The master array is

145
fixed on the high speed actuator which put the master array in circular motion. The axis of this circular

motion is parallel to the optical axis of the fiber, the radius is around 0.5[xni, that is relatively small

compared to the diameter of the fiber core. As shown in Fig. 2. by this circular motion, the center of the fiber core is also circulated about the

center 0. If the distance between 0 and the center of the fiber core of the sample array P is not so large

as the coupling coefficient approaches zero, the coupling coefficient shows a periodic change, the

frequency of which is equal to the frequency of the circular motion. A6 is the direction of the offset of the

fiber core of the sample array from O. and A6 is related to Acp which is the phase difference between the

periodic change of the coupling coefficient and the circular motion b}' the following equation:

AO = Acp (1)

Since the radius of circular motion r is relatively small compared to the radius of the mode field of the

fiber w. die mean coupling coefficient during the circular motion is almost equal to the coupling

coefficient in the case that the lateral misalignment is d which is the distance between 0 and P. So. is

expressed as followsf2|;

Tj = Tnexp

7'

4 +

(2)

(3)

where X is wave length, z is fiber separation distance, and w is mode field diameter. The range of the

periodic change of the coupling coefficient by the circular motion T,^ is gwQn by:

Tr = Td - I - Til + r

= Toexp

To

Tu exp

(4)

And the ratio of T^^ to T^, is related to d as follows: 7«

Tci = 2 exp

S c/ X —7'n r -/•- sinh '~2dr

11'" (5) ATui11'" exp

It means that d is approximateh in proportion to the ratio of T,^ to T^^. As it is clear from the above, in

regard to the offset of the fiber core, the direction can be obtained by measuring the phase difference Acp,

and the distance can be obtained b>' measuring the ratio of Tj^ to T^.

4. EXPERIMENTAL APPARATUS

Fig. 3 shows the system block diagram of measuring the core position. Single mode fiber arrays are

used in this experiment. The actaiator for the circular motion of the master array is a tube type piezo

actuator shown in Fig. 4. This actuator is specially made by changing the shape of the electrodes of the

piezo actuator that is used for the probe scanning mechanism of STM. By applying sme-wave voltage to

four electrodes around the tube and delaying each phase by nil. the master array fixed on the tip of the

actuator is given circular motion. In this experiment the radius is 0.5 |xm and the frequency is IkHz. In

146
Fig. 3. the light from LED. the wavelength of which is 1.31|xm. is incident mto the master array fibers,

transmits to the master array endface, radiates to the sample array. Part of the hght is coupled to the

sample arrav' fiber and detected by the photo diode. The phase difference between the detected signal

and the sine- wave voltage applied on the piezo actuator, and the amplitude of the periodic change of the

detected signal power are measured by two-phase lock-in amplifier. The mean power of the signal is measured by the voltmeter through the smoothing circuit. The sample array is fixed on the precision XYZ

stages and the displacement of the stages are measured by linear scale, the resolution of which are

0.01 Lun.

5. EXPERIMENT

For the confirmation of this method, the experiment to obtain the relation between the detected signal

at the photo diode and the displacement of the fiber core of the sample arra>' was made The clearance

between the master array and the sample array is adjusted at BOfajn. This clearance is filled with

matching oil and one of the fiber cores in the master arrav' is aligned to the fiber core in the sample array

by the XY stages with accuracy better than 0.05 [im. Then monitoring the lock-in amplifier and DC

amplifier, the X and Y stages are stepped b>' 0.1 [om. The displacement of the stages are measured by the

linear scales.

6. RESULTS

Table. 1 shows the relation between the phase difference Aqp and the direction of the offset of the fiber

core at the offset of Ijurn. The relative accuracy of the measurement is ±1°. Fig. 5 shows the relation

between Tj^/T^j and the distance. As equation (5) indicates.Tj^/Tj is in proportion to the distance . These

results confirm that by measuring Acp and Tj^/T^j. the direction and the distance of the offset of the fiber

core can be obtained. The repeatabilitv' of the measurement of Tj^/T^ is under 0.05 jim in the case that the

distance of the offset is around l[xm.

7. CONCLUSION

In this work, we investigated on the measurement method of the offset of the fiber core position of the

fiber array. Now we are developing the practical equipment for measuring the offset of fiber arrays using

this method. The cross-check between this method and conventional method is necessary.

8. REFERENCES

[1] M. Shigehara. "Investigation on Fiber Core Position Measurement Method for Multi-Fiber

Connectors", OFMC Proceedings. 1993, pp. 57 - 60.

[2] D. Marcuse. "Loss analysis of single-mode fiber splices", BellSyst. Tech. J., vol. 56, pp. 703-

718,1977

147
<—

tP"

a i

Figure 1 . Offset of core position

power

core of sample detected power

array fiber Tr

Td

time

core of master

array fiber

pliase of —-

circular motion Acp

Figure 2. Principle of measurement

LED 3

Piezo

Actuator

Master Array

Sample Array

XY Stages

Linear Scales

Sine-wave Generator

X

40mm

Lock-in Amp.

Low Pass Filter DC Amp.

E PD - Pre Amp.

Y+

Computer

Electrode

Figure. 3. System block diagram

Table 1 . Measured phase difference

14mm

Figure. 4. Piezo actuator

Ptiase Dif. <J(n=10) 0°(x:+1pm) fo°i 90°(y:+1pm) 90.6° 0.3° 180°(x:-1pm) 179.9° 0.4° 270°(y:-1|jm) 269.3° 0.4° 2.0

[pm]

0.0 0.5 1.0 1.5 Displacement of offset

Figure. 5. Relation between Tr/T^ and distance of offset

148
Polarization Mode Dispersion:

Definitions, Measurements and Statistics

N. Gisin

Group of Applied Physics

University of Geneva, 1211 Geneva 4, Switzerland

1 Introduction

Polarization Mode Dispersion (PMD) has the same effect on light signals in optical fibers as other kinds

of dispersion. In particular it limits the bandwidth and the linearity. It is caused by asymmetries in the geometry and/or the stress distribution in the fiber core. Locally PMD is due to the presence of

two group velocities, one for each polarization mode. Globally this effect is combined with random

polarization mode coupling. Consequently, PMD is a relatively subtle phenomenon.

In this contribution we like bring together several practical aspects of PMD: its definition, different measurement methods, and the statistical aspect of Polarization Mode (PM) delays. In the next section we summarize three measurements methods. A figure illustrates the setups and corresponding typical

results, and a table presents axivantages and drawbacks of each method. In section 3 two definitions of

PMD are proposed. They are related to each other and to the measurement methods. Some advantages

of each definition are discussed. Finally, in section 4 the probability that the instantaneous PM delay is larger than N times the mean PMD is computed and presented in a graph.

2 Measurement methods

The various measurement methods distinguish mainly by the coherence of the light sources. The polarimetric methods [1, 2, 3, 4] and the wavelength scanning method [5, 6, 7] all use high coherent

sources, that is light sources with a coherence time larger than the PM delay of the optical link under test. The reason is that all these methods require the light at the output of the link to be still polarized. The

way these methods analyze that outcoming polarization state varies, but they all require that state to be

stable during the measurement. Consequently, these quite elegant measurement methods have two main

limitations. First, they axe limited to relatively low PM delay (lower than the light source coherence time).

This difficulty is severe, since all these methods require a wavelength scanning that can only be realized

either with relatively expensive tunable lasers or at the cost of the dynamics of the measuring setup. Next,

the outcoming polarization state is highly sensitive to environment perturbations (vibrations, stresses, pressure, etc). Even a very local perturbation, on the centimeter scale, can deteriorate the measurement

of km long optical links. The interferometric measurement method [8, 9, 10, 11], on the contrary, uses a low coherent light source, like a LED. The outcoming light is thus depolarized and a local perturbation along the link has

only a limited effect on the measurement outcome, as for the measurements of other quantities, like the

attenuation for instance. This measurement is closely related to the direct measurement in the time

domain: it essentially measures the autocorrelation of the outcoming "light pulse" [12]. The coherence

time of the light source playing the role of the pulse width. This coherence time determines thus the

smallest measurable PM delay to about 0.15 ps. The figure 1 and 2 illustrate the basic measurement methods, typical setup and results, and their relative advantages and drawbax:ks. All other measurement methods can be considered as different technical implementations of these basic methods. In [13] the interferometric method is compared to the polarimetric method, and in [14] it is compared to the fixed analyzer method. In all cases the

nimierical results agree well then computed according to the same definition of PMD.

149
3 Definitions

There are essentially two possible definitions of an appropriate mean PMD.

The first definition follows closely the definition of modal dispersion for multimode fibers [15]. This

is natural since a "single" mode fiber actually supports several modes: the polarization modes. In the

absence of PM coupling, one has only two modes. But, in the more realistic case of random PM coupling,

the fiber behaves as a many-mode fiber and the natural definition for PM delay is the mean square

deviation of the time of flight of the various PM:

where I{t) is the intensity of the light pulse after propagating through the fiber, assuming an initial delta pulse. In the high PM coupling regime, I(t) is a gaussian of width Ar = D\JIjh where D is the

modal birefringence, h the mean PM coupling length and i the fiber length. The relation to the various

measurement methods is illustrated on figure 3. The second definition follows more closely optical communication applications where the laser source coherence time may be larger than the PM delay. In that case, for fixed PM couplings and a fixed

wavelength, each polarization state has a fixed group velocity. The fastest and slowest polarization states are called the principal states [1] . Note that these principal states, and the difference between their group

velocities, are very sensitive to the wavelength and to the PM couplings. However, it is natural to call instantaneous PMD the delay between the principal states and to consider their average. This leads us

to our second definition:

f^^ TUjS)djjj

PMDelayi t(cj)

'^'"^ (2) U2 — tJi We are now confronted to two "natural" definitions for one physical quantity! There is quite some

confusion in the literature about these definitions, and most comparative measurements have overlooked

the definition problem. However, in the cases of main practical interest (high and low PM coupling) these

two definitions are easy to relate:

For low PM coupling : PMDelay2 = PMDelayi

For high PM coupling : PMDelays = PMDelayi/ 1.08b (3) The "magic number" 1.085 ^/Stt/S) comes from the Maxwellian statistics of the delay between the

principal states [16, 17, 18] and from the following basic relation between measurement results with high

and low coherent sources [19, 13]:

PMDelayi = Ar = (4)

This relation is fundamental for the understanding of PMD, since it relates measurement results with

high and low coherent sources. Let us emphasize that this fundamental relation 4 is valid for any degree

of PM coupling. This is why we believe that the first definition, PMDelays, is preferable: it is directly

accessible by all measurement methods, without any assumption neither about PM coupling nor about

statistical distributions.

4 Statistics

In the high PM coupling regime, the statistics of the delay between the principal polarization states follows a Maxwell distribution [16, 17, 18, 13]. This is well established, as well from a theoretical point of

view, as from experimental evidence. Whether this regime is realized is a given fiber can be checked with

the polarimetric measurement method by directly investigating the statistics of the delay between the

principal states. It can also be controlled with the other two methods by comparing the interferogram

150
(or the fourier transform of the wavelength scanning data) to a gaussian. Indeed, the interferometric

method essentially measures the statistics of one component of the principal state vector. The statistics of these components is Gaussian whenever the statistics of the norm of the principal state vector (which

represents the delay) is Maxwellian (like the distribution of velocity of particles in a gas is Maxwellian

whenever the distribution of the velocity components is Gaussiaji). Figure 4, for instance, illustrates an

interferometric results in a case where the high PM coupling regime is not obtained. This figure should

be compared with the theory based on the telegraphers equation [9]. It clearly indicates that this fiber

is inbetween the low and high PM coupling regime.

The high PM coupling regime is the most relevant one for practical applications, since the other regime

appear only in special fibers or in fibers with large geometrical defects. In this case the probability that

the instantaneous PMD is larger than N times the mean PMD At (1** definition) is given by:

Probability{PMD > N * At) ^ / —== exp{ 7-r^)dT (5) Jn*At y/n/2 (Ar)3 2(At)^

and is illustrated in figure 5.

5 Conclusion

Polarization Mode Dispersion is an important parameter for optical communication, as well for CATV

(linearity [20]) as for long distances. The actual PMD of installed cables varies greatly, even from one

fiber to another in the same cable. Typical values range from lower than 0.lps/\/km to Ips/ \/km, with

exceptional values as high as Aps/ \/km [21, 22]. Reliable PMD measurement are nowaday feasible in a routine way. However, attention should be

paid to the definition of the mean PMD and to the statistics of the instantaneous value of PMD.

6 Acknowledgments

Financial support from the swiss "Office Federal de I'Education et de la Science" within the COST 241 european

project is gratefully acknowledged.

7 References

1. CD. Poole and R.E. Wagner, El. Lett. 22, 1029 (1986).

2. B.L. Heffner, IEEE Photon. Tech. Lett., 4, 1066, 1992.

3. D. Andresciani et al. Optics Lett. 12, 844, 1987.

4. N.S. Bergano, CD. Poole and R.E. Wagner, J. Lightwave Tech. 5, 1618, 1987.

5. CD. Poole, Optics Lett. 14, 523, 1989.

6. L. Thevenaz and Ph. Robert, ECOC'90, 1990.

7. CD. Poole and D.L. Favin, "PMD measurements based on transmission spectra through a polarizer", J. Lightwave

Tech., in press, 1994.

8. N. Gisin, J. P. Pellaux and J. P. Von Der Weid, Symposium on Optical Fiber Measurements, pp 119-122, 1990.

9. N. Gisin, J.P. Pellaux and J.P. Von Der Weid, J. Lightwave Tech. 9, 821, 1991.

10. Y. Namihira et al., J. lightwave Tech. 1, 329, 1989.

11. Y. Namihira et al., J. Optical Communication 12, 1-8, 1991.

12. Let us emphasize that the spread A (ie the standard deviation) of an autocorrelation function is always ^2 the spread of the initial function /: A(/ * /) = V^A(/)

13. N.Gisin, R. Passy, JC BischoflF and B. Perny, IEEE Photon. Tech. Lett., 5, 819, 1993.

14. N.Gisin, R. Passy and J.P. Von der Weid, Photon. Tech. Lett., "Definitions and Measurements of Polarization Mode Dispersion: Interferometric versus fixed analyzer methods", in press, June 1994.

15. CCITT Recommendation G.651 - Characteristics of a 50/126 micron graded index optical fibre cable. COM.

XV-R-89 E, May 1992.

16. F. Curti, B. Daino, G. de Marchis and F. Matera, IEEE J. Lightwave Tech. 8, 1162 (1990).

17. N. Gisin, Opt. Communications, 86, 371, 1991.

18. G.J. Foschini and CD. Poole, J. Lightwave Tech. 9, 1439, 1991.

19. N. Gisin and JP. Pellaux, Optics Commun. 89, 316, 1992.

151
20. CD. Poole and T.E. Darcie, J. Lightwave Tech. 11, 1749, 1993.

21. N. Gisin, B. Perny and R. Passy, Proc. Optical Fiber Measurement Conference, York, 1991.

22. N. Gisin, OFMC'93, Torino, 1993.

PMO Graph - Mes:MA-A012 - ZoomMOCW.

Delay [plco seconds] Wavelength : 1.55 ( m]

Fixed Analyser measuremeni

1. Typical measurement setups and results for the three basic measurement methods for PMD. a) interferometric,

b) wavelength scanning, c) polarimetric

152
153
CD

154
MODE COUPLING MODEL FOR THE JONES MATRIX

EIGENANALYSIS METHOD

D. A. NOLAN AND D. Q. CHOWDHURY

CORNING INC., SULLIVAN PARK, CORNING N.Y. 14831

TEL 607 974 9000, FAX 607 974 3675

Abstract

A simple model is proposed for estimating an effective fiber

birefringence and mode coupling length from the wavelength dependence

of a fiber's polarization mode dispersion (PMD). The model is

particularly well suited for use with the Jones Matrix Eigenanalysi

s

( JME)

.

I. Introduction

A number of measurement techniques have been proposed for measuring

PMD. The Jones matrix eigenanaly ;5is method is well suited for

determining this parameter. With this technique, one can determine

the wavelength dependence of PMD as well as a mean PMD value over a range of 100 nms . We present a model whereby one can use this data to

determine an effective mode coupling length and also the effective

intrinsic fiber birefringence. In Section II, we develop the model,

and the parameters of interest are derived. As an example, we then

calculate these parameters from the wavelength dependent data of a dispersion shifted fiber with a moderate amount of PMD.

II . Model

Fiber-polarization mode dispersion is modeled using a concatenation

of low birefringence polarization maintaining fibers and wavelength

independent couplers. Fig. 1. The' PM fibers simulate, for example,

the effect that core deformation has on polarization dispersion. The

fibers are chosen to be of equal length. This simplification is not

necessary, but at this point is imposed for the purpose of keeping

the model simple. The validity of this assumption can be checked as

discussed below. The field amplitudes after propagating through n - 1

couplers can be expressed as

155
N-1 |a. b. I |exp[-joo8T. ] 0 1 | v|yo^ | n I ^ 'I I '

i I ^ I i=i I|b. a.|

I I III I 0 1| I ii/Oj^ I

( 1

)

where a^ and b^ are the mode coupling and are estimated as

a. = /(I - a.)

bi = j/(a.) (2)

In eq. (2), the coupling coefficients ( ) are chosen to be .5 in

order to impose a square root length dependence on the link PMD. From

the field amplitudes, the time delay differences can be calculated

1 using the Jones matrix technique'

1^=1 I exp[-:w5T ] 0|

I ^1 = I ^ I III I l^bl I 0 1 1

At = I Arg( pl/p2 )/Aco| ( 3 )

In eq. (3), and p2 are the eigenvalues as determined in Ref. 1,

and Aco is a difference in optica], frequency. Using this equation, we

have calculated the wavelength dependence of the PMD for a number of

different model parameters. The frequency of the oscillations in all

cases is directly related to the value of St., as was predicted also

in Ref. 2.

Similar to the predictions of our model. Fig. 2 shows the dependence

of PMD as measured on 2 kilometers of dispersion shifted fiber. Based

on our model, one can determine the &t^ as measured from the maximum

of the Fourier transform or other power spectral density methods.

P(&T^) = l./2ii J PMD( w)exp[ j(o6T^ ]doo (4)

The value of 6t. can also be estimated by picking wavelength

differences at which a phase change of k/4 occurs and using the

expression

2jic[1./X^ - I./X2] St. - ti/2 .

.- (5)

156
For this fiber and using these techniques, we have determined to

be .02 psecs. The effective mode coupling length, 1^, is now

determined by imposing a square root length dependence on the

polarization dispersion and using ( see Ref. 3, for example )

<PMD>L = St. [l./L] -5 - (6)

In eq. (6), <PMD> represents the wavelength average of the link PMD

.

Using .2 psecs for the link PMD, 2 Kms for the link length, and

.02 psecs for &t^, we determine the 1^ to be 40 meters for this

particular fiber. The assumption of similar St^ s throughout the

fiber is valid if the fourier transform of eq. (4) has only one

strong maximum. This fiber did show only one strong maximum. In a subsequent publication, we will consider the case of variable St. s.

In summary, a mathematical model used in conjunction with the Jones

matrix eigenanalysis technique can be used to interpret fiber mode

coupling parameters, i.e. coupling length and effective

birefringence, St^/I^.

REFERENCES

1. B. L. Heffner, 'Automated Measurement of Polarization Mode

Dispersion using Jones Matrix Eigenanalysis'. PTL A, 930-932,

1992.

2. N. Grisin and J. P. Pellaux, 'Polarization Mode Dispersion: Time

Versus Frequency Domains', Opt. Comm. 8_9, 316-323 , 1992 .

3. D. A. Nolan, R. M. Hawk, and D. B. Keck, 'Multimode Concatenation

Modal Group Analysis, JLT 5, 1727-1732, 1987.

157
Slow Axis St. fix. 5x.

Fast Axis n \ ^ V V \^ A A /V

• I are lengths that introduce PMD through core

ovality

• C. are couplers simulating mode coupling

through bends and twists

Fig. 2: Wavelength vs. PMD

158
Study on A Conservation Quantity in PMD Measurements

Osamu Aso, Isamu Ohshima and Haruki Ogoshi

Funikawa Electric Co., Ltd.

6, Yawata Kaigandori, Ichihara, Chiba 290, Japan

1. Introduction

Recent progress of optical amplifier and chromatic dispersion compensation technique makes it possible to

realize high-bit rate and long haul repeaterless communication system. When chromatic dispersion is completely

negligible, PMD will limit the system performance. Then suitable PMD compensation technique is necessary.

In spite of the above, importance of PMD compensation, random mode coupling and its time evolution makes it

difficult to estimate PMD of single-mode fibers sufficiendy.

In this paper, through the evaluation of the critical wavelength step-size'"TM using 10 - 90km dispersion shifted

single-mode optical fibers ( DSF ), we will study the accurate measurement technique of the Jones matrix

eigenanalysis.

In Sec. 2 for the convenience of the reader, we give a brief summary of the significance of cridcal wavelength

step-size AXc. In Sec. 3 a novel method to determine the zlAc is introduced. In Sec. 4 experimental results are

shown. Finally, conclusion and discussion are given in Sec. 5.

2. Critical wavelength step-size AXc

According to Jones matrix eigenanalysisTM, in order to estimate the group delay difference of principal states at a

wavelength A, it is necessary to measure Jones matrix over the wavelength interval AX , i.e., at A and X + AX.

As shown in ref [IJand [3], opdmum wavelength step-size should be narrow enough to satisfy exponential

approximauon but wide enough to provide the best accuracy ( the wider the step-size the better the accuracy ).

Recently the authors confirmed experimentally'^' , using an about 90k;m DSF, that the exponendal

approximation region almost coincides with the Poole's first order approximation region^'. Thus hereafter we

will use the word " first order " approximation to be synonym with " exponential " approximation.

The above discussions indicates the existence of the transidon point AXc\ the critical value of transition from

the first order dominant region to the higher order region. Obviously AXc depends the type and magnitude of

birefringence, thus estimadon of this value will give an important information on accurate PMD measurements.

In all the following discussion, we will call the above opumum step-size as cridcal step-size AXc.

3. Experiments

We used Jones matrix eigenanalysis described in Sec. 2 and ref [3] to measure PMD of DSFs. Light from a A =

1550nm ( zero-dispersion wavelength ) and X + AX ait passed through the test fibers and measured degree of

polarization was greater Uian ~ 95% after traversing test fibers. All of the DSFs used in our experiments were

fragments of a 90km optical fiber. Fibers were wound onto 280mm diameter drum at 20gw tension. In order to

evaluate whedier a wavelength step-size AX is inside the first order approximauon region or not, we consider the

global and local measurement ( see Fig. 1 ).

First of all PMD is measured over the wavelength interval { X, X + AX). This measurement is called hereafter

159
"global " measurement and its result is described as zi t. Then AX is divided into AX / 10 intervals and PMD is

measured over these intervals respectively. This is called " local " measurement and each result is Av ( / = 1, 2,

, 10 ). When AX is larger than AXc, higher order effect is not negligible and principal states will be violated

over the wavelength interval. Thus, in order to evaluate whether the step-size AX allows first order approximation or not, we consider following ansatz for the satisfying condition of the first order region.

10

1 < cr (1 ) where, < Ax> and o is averaged value and standard deviation of the 10 times global measurements. In Fig. 2, we will show the comparison between global and local measurements in the case of SOkm DSF. From this

result, critical wavelength step-size is estimated as zi Ac = 6.0 — 1.0 nm.

4. Results

Fig. 3 and Fig. 4 show the length dependence of PMD and critical wavelength step-size respectively. Relation

between PMD and fiber length L is suitably approximated Ax I = 0.048 psec. km and also the relation

between critical wavelength and fiber length seems to be obtained zi/L L'" = 42.8 nm km"^

The above discussions are confirmed to by considering the following quantity A : At AX c _ ( 2 ) As shown in Fig. 5, A is almost independent of fiber length L and its value is estimated as y\ = 2.05 psec. nm.

5. Conclusion and discussion

Optimum wavelength step-size was discussed experimentally as a critical step-size due to the transition from

first order dominant region. A practical method to determine A Xz was obtained and length dependence of zi was

estimated. In our result, AXc \s proportional to L and this result implies the existence of a conservation ( length independent ) quantity. Hence we proposed yi-term, which is defined as the product of the critical

wavelength step-size and PMD value measured by this step-size. As shown in Fig. 5, this quantity is almost

independent of fiber length.

Finally, we consider the appUcation of y\-term. As shown in Fig. 6, we plotted the curve Ax AX = A in Fig. 2.

This curve pass through the transition point and we obtained almost same results from other measurements with

various length. Thus A-term could hopefully be a tool to confirm PMD measurement accuracy. Of course this is

a result of our experiments. Then to confirm reproducibility of this result is necessary.

References

[1] Hewlett Packard Co., " HP 8509 Lightwave Polarization Analyzer, Technical Data ", ( 1993 ).

[2] O. Aso, I. Oshima and H. Ogoshi, " A Study on The First Order Effect of The Phenomenological PMD

Model in Long DSF. ", in Press ( 1994 ).

[3] B. L. Heffner, " Automated Measurement of Polarization Mode Dispersion Using Jones Matrix

Eigenanalysis. ", IEEE Photon. Technol., Lett., vol. 4, No. 9 ( 1992 ) 1066 - 1069.

[4] C. D. Poole andR. E. Wagner, " Phenomenological Approach lo Polarization Dispersion in Long SingleMode Fibres ", Electron. Lett., vol. 22, No. 19 ( 1986 ) 1029 - 1030.

160
At

global measurements

<

X -[ \ \

1

local measurements ^^^^

Fig. 1 Schematic diagram of global and local measurements.

0.40

<1

0.10

: * 5 J 2 5 iQooioooooooooi

- "

I

0 local measurements

global measurements

0.0 5.0 10.0 15.0 20.0 Ak ( nm )

Fig. 2 Comparison of global and local PMD measurements. # and O indicates < 4t> and averaged

value of local measurements respectively. Error bar represents standard deviation.

1.00

o

0)

(/) CL 0.10

0.01

1.0 10.0

L ( km

100.0

Fig. 3 Plot of PMD value as a function of fiber length. Error bar indicates standard deviation from

10 times measurements. Dashed line represents fitting formula: 4t/L"' = 0.048 psec./ km"^.

161
T f

( 1 1 1 1 1 1

1

0.0 20.0 40.0 60.0 80.0 100.0

L ( km)

Fig. 4 Plot of critical wavelength step-size as a function of fiber length. Error bar indicates standard deviation from 10 times measurements. Dashed line represents fitting formula: A/icL'" =42.8 psec. km'

40.0 60.0

L ( km )

100.0

Fig. 5 Plot of /\ term as a function of fiber length. Error bar indicates standard deviation from 10 times measurements. Dashed line represents fitting formula: A = 2.05 psec. km"^.
Quantitative comparison between the Jones matrix

eigenanalysis & the wavelength scan method for

polarization mode dispersion measurement: a

simulation study.

D. Q. Chowdhury and D. A. Nolan

Corning Inc., Sullivan Park, Corning, NY 14831

Phone: (607) 974-3593 Fax: (607) 974-3675

Abstract

A model representing a long optical fiber by a random concatenation of small linearly birefringent fibers is used to simulate the measurement of the polarization mode dispersion (PMD)

with the use of the Jones matrix eigenanalysis (JME) method and the wavelength scanning

(WS) method. Simulations indicate that for the WS method the ensemble average power

spectral density (PSD) is a Gaussian and the final ensemble average PMD values agrees well

for a large number of random samples. However, PMD values obtained with the use of a

single measurement of both methods could differ significantly.

1 Introduction

There exists several methods for measuring PMD of a long optical fiber. AU of them can be categorized under three fundamental methods: (1) Jones matrix eigenanalysis; (2) wavelength scanning technique; and (3) interferometric technique. Since there is no industry

standard on the PMD measurement technique, it is of interest to compare each of the methods quantitatively. However, the fact that the PMD of a given fiber is a statistical variable

has made such comparisons with actual measurement very time consuming and impractical.

The statistical variation of PMD depends on both the inherent fiber perturbations built-in

during fabrication and the environmental perturbations from the deployment conditions.

Consequently, numerical simulation is one of the easier ways to make such a quantitative

comparison between different measurement techniques.

In this paper we compare two of the above three techniques on a quantitative basis using

computer simulation. First, a brief description of the theory of simulation is presented. Then

some quantitative comparison between the JME method and the WS method are presented.

Using a Monte-Carlo simulation we show that the ensemble average Fourier transform of the

WS method approaches a Gaussian as the number of random samples increases.

2 Model

A model to simulate the PMD measurement with the use of the JME and the WS method is developed based on the derivations in Ref. 1. The concatenation model is sHghtly altered in

163
order to model the WS measurement. The PMD at the end of the N^"^ link of a concatenated

fiber is given by [2] where = 7JveAr+(m^7^-e^)e;v+cos(7jva;)[m^7^-(m^7^-ejv)eAr]-sin(7^a;)e7vXm^7^ (1)

In the above Eq. is a vector representing the principal states of polarization (PSP) at the output of the N^^ link on the Poincare sphere, cat is the PSP vector for the link iV, 7jv

is the total mode group delay between the the two PSP's of Hnk N, and uj is the angular

frequency of the incident light.

In order to simulate the WS method, we use an i-polarized incident wave propagating

through the randomly concatenated fiber Hnks. The output state of polarization (SOP) is given by the recursive relation = (^^7^*eiv)eiv + cos(7jvw)[n^7i - (n^7^-ejv)ejv] - sin(7jvu;)eiv Xn^7\ (2)

where is the vector indicating the output SOP at the end of the N^^ hnk on the Poincare

sphere. The polarization of the detected hght is also assumed to be x-polarized. Since the

concatenation is random, the direction of the detected polarization reaUy does not matter.

The intensity of the detected polarized hght is computed as .

/x= = 0.5(1 + nl-e,), (3)

where Ej. is the unit vector along the x-direction in the Poincare representation. Note that

is a vector representing the PSP and could be greater/less than unity in magnitude.

Whereas, is a vector representing the SOP with a magnitude of unity for a lossless fiber.

3 Processing the WS data

There are two different ways to obtain the PMD value from the WS data. One of them uses the number of peaks in the WS data within a given wavelength interval. The other method

fits the power spectral density (PSD) of the wavelength scan with a Gaussian of the form

exp[—r^/2<T^] and the PMD is given by the average deviation o-y^/TT. For the WS method

we use the latter of the techniques to obtain the PMD value.

In order to obtain a PMD value from the WS measurement we use a fast Fourier transfrom

(FFT) algorithm [3] to obtain the PSD which is then fit with a Gaussian function. For the

FFT algorithm the polarized intensity vs. frequency data is already normalized in our model

since the fiber is assumed lossless. The computed data is multipHed by a Hanning window [3]. Mathematically expressed

R{i) ^ R{i) \[l - cos {j^)] , (4)

where R(i) is the FFT data set, J is the total number of data points in the FFT. We subtract

the average of the wavelength scan over the scanned range from the data to make the average

zero and thus remove the zero-frequency contribution.

164
4 Results and Discussion

In our simulation the number of concatenated links, N, representing a long optical fiber is computed from the fiber length and the average Unk-length. Since the PMD of a given fiber

is statistical in nature, in order to compare PMDs measured with the use of JME and WS

methods, a large number of random environment (deployment) samples are required. In our

simulation a random number generator is used to vary the individual Hnk lengths about the

average along with their PSP orientations. For a quantitative comparison 400 random seeds

are used in the simulation to compute the average PMD using both methods. The average

and the standard deviation of the PMD for the JME simulation is computed using aU the

wavelength values for all the random samples.

< Ar>jME <Ar>ws cTvvs/ <Ar>ws

0.0290 0.0123 0.4223 0.0296 0.0123 0.4171

0.0394 0.0166 0.4223 0.0370 0.0170 0.4596

0.2522 0.1090 0.4322 0.2078 0.0744 0.3582

0.3585 0.1592 0.4442 0.2976 0.1196 0.4020

0.5676 0.2474 0.4359 0.4373 0.1546 0.3536

0.7865 0.3407 0.4332 0.6332 0.2519 0.3978

Table 1: Comparison between the JME and the WS method measurement simulations. <Ar>'s

are the average PMDs and cr's are the corresponding standard deviations of the average PMDs.

400 random samples are used. The subscripts indicate the method simulated. Fiber length was

1 km composed of 20 Links with an average link-length of 50 m. The PMD values in the table are

in ps/km.

Table 1 shows that for a low PMD fiber the average PMD and the standard deviation

of the average agree well. Also note that in the JME simulation, even for a low PMD,

the ratio of the standard deviation to the average is much closer to the same ratio for a MaxweUian distribution (0.422). For larger PMDs both the simulated averages and the

standard deviations for the WS method differ from those for the JME method. Also for the

WS simulation the ratio between the standard deviation and the average differ from 0.422.

This is also true for the WS method for a long fiber with a low per- unit-length PMD.

The typical WS instrument currently has a precision of about 2 nm in wavelength, in

other words, AA could be as small as 2 nm. This AA corresponds to a Fourier transformed

range of 1.6 ps 27rcAA/A^) within the typical available wavelength range (1150 - 1700 nm).

The accuracy of the Gaussian fit should start to deteriorate at PMDs greater than 1.0 ps. Because of the statistical nature of the PMD, we decided to look at the ensemble average

of the FFT taken over a lan^ number (200) of random samples. Figure 1 shows the average

FFT of 1, 6, and 200 random samples generated by the computer. Power spectral densities

(PSD) computed for a single random sample using the maximum entropy method (MEM) [3] are also shown in tl^^gure.

Figure 1 shows thelt as the number of random samples in the ensemble average increases,

the FFT conveiges to a Gaussian. Even with as few as 6 random samples the average FFT

165
t

- — Gaussian fit •••• Fi=Tof 1 — <FFT> of 6 <FFT> of 200

0.4 0.6

time (ps)

MEM PSD

Gaussian fit

0.0 0.2 0.4 0.6 0.8 1.0

time (ps)

Figure 1: The average FFT for 1, 6, and 200 random samples along with the Gaussian fit for the

average FFT of 200 random samples. The power spectral density (PSD) computed with the use of

the maximum entropy method (MEM) and the corresponding Gaussian fit is also shown. The PSD

is estimated for a single wavelength scan for which the FFT is also shown. Average PMD obtained

from the average FFT of 200 samples is 0.28 ps, and the average PMD obtained from the single

MEM PSD is 0.23 ps. Average PMD for a single FFT is 0.35 ps.

approaches a Gaussian distribution. Figure 1 also indicates that to estimate the PMD from

a Gaussian fit of a single WS data could be erroneous. However, simulations (not shown

here) suggests that the MEM could provide a more stable solution to this problem. Unlike

FFT, MEM does not assume that the WS data is periodic after the measured wavelength

range and, hence, provides a more realistic estimate of the PSD.

5 Conclusions

With the use of computer simulations we have shown that when an ensemble average is considered, the JME and the WS method agree well. However, for a single measurement,

the WS measurement could differ significantly from that of the JME measurement. The

fact that the ensemble average FFT is Gaussian gives more credibility to the FFT method

than the "peak-counting" method, also used to obtain PMD values from WS data. However,

PMD value obtained from a Gaussian fit to the FFT of a single WS data could be erroneous. A possible solution could be to use a more realistic PSD estimator, for example, MEM. The

largest PMD that can be measured with the use of the WS method and FFT is limited by

the resolution of the wavelength increment. This could be a Limiting factor if a soHd-state

monochromator is used to spectrally resolve the wavelength vs. relative amplitude data.

References

[1] N. Grisin and J. P. Pellaux, "Polarization mode dispersion: time versus frequency domains," Opt. Comm. 89, 316-323 (1992).

[2] Equation 3 in Ref. [1].

[3] W. H. Press, B. P. Flannery, S. A. Teukolsky, and W. T. Vetterling, Numerical Recipes,

The Art of Scientific Computing (Cambridge University Press, Cambridge, 1989).

166
POLARISATION MODE DISPERSION MEASUREMENT USING FOURIER ANALYSIS:

INVESTIGATION OF MODE COUPLING BEHAVIOUR IN FIBERS

A J Barlow, EG&G Fiber Optics, Wokingham, U.K.

Introduction

Several techniques for the measurement of polarisation mode

dispersion (PMD) have been proposed, and recently compared [1].

Many of these methods operate in the frequency domain, relying on

relatively simple instrumentation and data collection, and Jones

or Stokes analysis methods to predict the PMD. Conceptually,

however, there are some difficulties [2] in using the matrix

techniques to relate to the real picture of pulse broadening in

the fiber, which is the root cause of PMD-related bit errors or

analog distortions in optical communications systems. Methods

which display this data [1], [3] succinctly have great appeal.

Another major factor is that PMD in communications fiber is

inherently statistical, and any measurement approach must take

this into account. PMD is not a single-valued parameter, and the

full distribution of pulse arrival times is required. Measurements are often required in the final operating environment to

characterise the actual fiber deployment. The Fourier transform

technique used here is inherently statistical in approach and

combines the twin advantages of data collection in the frequency

domain, and a succinct visual data output of the pulse arrival

time distribution. In this paper, this method is applied to

several fibers to investigate their mode coupling behaviour.

Background

The well-known cycle counting (spectral scanning) or fixed

analyser method is essentially a polarimetric (frequency domain)

method, where the light transmission thorough the fiber between

randomly oriented polarisers is monitored as a function of wavelength, relative to a scan without the output polariser (analyser) [3]. A typical spectral scanning transmission plot R(X) for

a 4.4km standard fiber on a shipment spool is shown in Figure 1.

The periodicity of this waveform is determined from the minima

and maxima of the waveform over a given wavelength interval, and

the PMD found from the number of cycles in the waveform. Moreover, because it uses a very large spectral range, it is possible

to obtain a good spectral average or "ensemble" of the fiber mode

coupling effects, and thus a reasonable prediction of the total

fiber PMD [1], and moreover with relatively simple instrumentation.

Nevertheless, considerable additional information can be obtained from the data in figure 1, by making use of the obvious

relationship between the frequency domain where the data is

collected and the time domain in which we require our PMD output, namely the Fourier Transform. A Fourier transform on R(X)

transforms the optical frequency domain data to the time domain

yielding directly the light arrival time distribution P(6t;) and

167
RMS pulse spreading <£\,t> [2]. The Fourier method is applicable

to all fibers with weak or strong mode coupling as well a fiber

components eg isolators.

To briefly show this relationship, the fiber under test between

the polariser and analyser/polarimeter is modelled by a birefringent element placed between crossed polarisers, each at +/-

45° to the fiber birefringence axes. The light transmission ratio

R as a function of optical frequency i> = cq/A given by:

R = sin2 (n. {i\n. L/cq} . ) (1)

where

CQ is the velocity of light in vacuo.

i\n is the linear birefringence of the fiber per unit

length,

L is the fiber length

Now, the PMD of the device is defined as:-

6t: = An.L/CQ (2)

where we have neglected any (usually small) dispersion term in (\n

for clarity. Using equations 1 and 2 we obtain a cosine wave of

oscillation frequency of 6t;. The spectral content of R obtained

by Fourier transformation would exhibit a single 'spike', at 6t;,

so the PMD is the centroid of this spike.

For the case where the fiber has strong mode coupling, we can

see that the individual coupled 'paths' will arrive at specific

relative times, giving rise to a corresponding periodicity in

R(X) and hence a 'spike' in the Fourier transform distribution,

with an amplitude corresponding to the relative power being

carried. The actual Fourier Transform obtained is a "folded"

symmetrical one, since negative 6t values appear as positive.

Random coupling leads to a Gaussian distribution; non-random

coupling to an undefined distribution pattern. In either instance, the second moment, Opj, of the distribution P(6i), is

related to the expected value of PMD <|iT> - the RMS pulse broadening - of the fiber. The Fourier method offers extremely rapid

and statistically accurate characterisation of the PMD and the

distribution within the broadened output pulse. By chi-squared

testing of the distribution P(6-c), we can determine within

statistical limits of uncertainty whether the data 'sample' taken

during a measurement is Gaussian and therefore indicative of

random mode coupling in the fiber [3].

Results

An example of the Fourier transform output obtained in a 25km

long fiber on a normal shipment spool is shown in Figure 2. In

this case the data is indeed Gaussian (random) as expected for

such a long fiber with multiple layered windings. Experience

shows that a great many such fibers are mainly Gaussian, provided

they are greater than ~10km in length. Spooled fibers shorter

than this can sometimes be non-random, and an example is given in

Figure 3, which is the FFT plot for the 4.4km fiber data of

Figure 1. A prominent lobe in the distribution is clearly visible, and appears to correspond to the bend birefringence at the

spool diameter. The bend birefringence appears to provide suffi168
cient deterministic birefringence behaviour to dominate other

sources of mode coupling (eg twist) so the fiber starts to behave

non-randomly . Here of course, the PMD no longer builds up with a

root- length dependence. Only when the fiber length and number of

winding layers build up, does sufficient mode coupling apparently

occur to randomise the effect of the bend. Further investigation

and results will be presented at the Conference in order to

confirm this.

Discussion

It is clear from these examples, that the PMD pulse broadening

can vary enormously from fiber to fiber, and further study of

these effects is warranted. Given that the pulse distribution

data can visualised in any fiber by the Fourier method, further

study using this method has considerable merit. Because of its

statistical nature, PMD data must necessarily be gathered over

many fiber samples in order to validate any conclusions of such a

study.

As shown above, the Fourier method is able to provide the full

pulse broadening distribution data in a matter of minutes in both

lab and field conditions. The short measurement time allows

routine field and/ or production measurements of PMD to be contemplated, in order to build up sufficient fiber samples for an

adequate study. Of significant concern is that fiber deployment

conditions, eg measurement drum, shipment spool, cable, installed

cable, will significantly affect the PMD that results.

Conclusion

PMD measurements using the Fourier Analysis (Spectral Scanning)

technique have been described. The method offers the advantage

of relatively low cost simple instrumentation, coupled to direct

visual output of the broadened output pulse resulting from PMD

and mode coupling.

The method has been applied to investigate the mode coupling

conditions prevailing in typical fibers, relating these to the

deployment and fiber length.

Finally, by enabling the user to quickly and easily examine the

pulse power distribution, the Fourier method offers significant

potential for routine production and field PMD measurement, which

will also contribute to the understanding of PMD and mode coupling in fibers under a variety of production and deployment

conditions. A full understanding of these effects is essential to

the reliable control of PMD levels in installed fiber links,

based on simple rapid production screening of PMD.

References

1) . N. Gisin, "Definition of Polarisation Mode Dispersion", to

be published.

2) . B Heffner, "Polarisation Mode Dispersion", Tutorial paper

ThE, OFC 94.

3) . TIA FOTP-113 "Polarisation Mode Dispersion Measurement for

Single-mode Fibers by Wavelength Scanning", draft, FO. 6.6.5.

169
Figure 1 : Ratio Plot obtained

for 4l<in fiber on siiipment

spool.

1500 1550 Wavelength (nm)

0,05-

0.045 • 004

0,035- I

0O3- S 0025-

I

0,02- a

0.015-

0 01-

0.005-

Figure 2: Fourier Transform

for 25km fiber on shipment

spool. Data is GAUSSIAN

0 05 0.1 0 15 0 2 0 25 0 3 0 35 0 4 0 45 0 5 Relative Pul9e Arrivol Time (psf

FFT Magnitude -A- Gaussian fit

Rgure 3: Fourier Transform

for 4km fiber on shipment

spool, corresponding

to the ratio plot In Figure 1. DataisNON GAUSSIAN

Relative Pulse Arrival Tin^ (ps)

170
INTER-LABORATORY POLARIZATION

MODE DISPERSION MEASUREMENT STUDY

W. B. Gardner, TIA Ad Hoc Group*

AT&T Bell Laboratories

Norcross, GA 30071

*A. J. Barlow (EG&G Fiber Optics);

K. A. Emig (Coming); Scott Grindstaff (AT&T Fitel);

Paul Hemday (Hewlett-Packard); W. C. Jones (Bellcore);

Tom Olson (Alcatel); D. K. Roland (SpecTran).

1. INTRODUCTION

Polarization Mode Dispersion (PMD) causes optical pulse spreading which could impair the

performance of systems. PMD also causes the state of polarization (SOP) to vary with

wavelength, and this latter effect provides a convenient means for measuring PMD. One version

of this is the wavelength scanning method\ sometimes referred to as the "fixed analyzer"

method^. Three ways of implementing wavelength scanning are illustrated in Fig. 1. Each relies on measuring how rapidly the SOP emerging from the test fiber changes as the wavelength is scanned. This paper reports the results of an inter-laboratory comparison of PMD measurements

using wavelength scanning, initiated by TIA's FO-6.6.5 Working Group on Single-Mode Fiber.

2. BACKGROUND

Light from a narrow-band source will encounter a differential group delay time Ax between the

two orthogonally polarized principal states^ of polarization. A long fiber produces sufficient

random coupling between polarization modes to cause PMD to be a stochastic variable. PMD

is defined as <Ax>, where < > denotes the average. The average can be taken over time,

temperature, or wavelength. The wavelength scanning method used in this study averages over

wavelength. The differential delay time Ax causes spreading of digital pulses in the time domain.

In the limit of strong mode coupling, <Ax> is directly related to the rms pulse broadening. Fibers

in practical systems are nearly always in this regime, and in that case the PMD coefficient is defined as <Ax>/L"^, where L is the fiber length.

3. THE MEASUREMENT METHOD

In versions (a) and (b) of wavelength scanning shown in Fig. 1, the transmitted power P(X) is measured as a function of wavelength. This measurement is repeated with the Analyzer removed

from the beam. The resulting power, P„(k), can be used to eliminate the spectral dependence of

the measurement system components and the test fiber loss. Typical plots of the ratio R(k) =

171
P(k)/P^(X) are shown in Fig. 2. Version (c) produces three normalized Stokes parameter

responses from a single measurement.

If R(k) has N maxima (or minima) within the window Xi<X<k2, then

<At> =

(1)

where c is the speed of light in vacuo, k is a mode coupling factor which equals 1.0 in the

absence of mode coupling, and 0.82 in the limit of strong mode coupling"*. This factor accounts

for the fact that the direction of the polarization dispersion vector on the Poincare sphere varies

with frequency in the presence of mode coupling. The PMD in Eq. (1) is interpreted as an

average over the wavelength range Xi<X<^2-

Algorithms for identifying peaks in R(A,) have been developed. Power spectral density estimation

techniques, such as the Fourier Transform, are an alternative to cycle counting. Participants A

through G used peak counting, while H used Fourier Analysis^

4. RESULTS AND DISCUSSION

Two depressed-cladding MCVD single-mode fibers, each about 5 km in length, were wound

under 30 grams tension onto 15 cm dia. spools. PMD values from Eq. (1) were divided by the

square root of the fiber length, and are listed in Table 1 in the order in which the measurements

were made. Since the purpose of the study was to compare test sets, k in Eq. (1) was taken to be 1.0 for all measurements.

Although the standard deviations (a) in Table 1 may appear to be large, the biggest contributor

to a is in fact the random fluctuations in Ax. This contribution to a is given by Eq. (38) in Ref.

4:

This contribution, which appears in the first row of Table 2, is inherent in the nature of PMD,

and would be present even with perfect test sets.

5. CONCLUSIONS

Eight participants used the wavelength scanning method to measure the PMD of two 5 km fibers.

The statistical uncertainty inherent in PMD (about 20% for the fibers and wavelength ranges used

in this study) is the major contributor to the observed inter-laboratory variation.

a2/<AT>2 ^ O.llX^X^/ [<At>(A2-Ai)c] (2)

172
References

1. CD. Poole, "Measurement of PMD in Single-Mode Fibers with Random Mode Coupling,"

Optics Lett. 14, 523 (1989).

2. Y. Namihira and J. Maeda, "Polarization Mode Dispersion Measurements in Optical Fibers,"

NIST Symp. on Opt. Fib. Measurements, Boulder, September 15-17, 1992, p. 145.

3. C. D. Poole and R. E. Wagner, "Phenomenological Approach to Polarization Dispersion in

Long Single-Mode Fibers," Electronics Lett. 22, 1029 (1986).

4. C. D. Poole and D. L. Favin, "PMD Measurements Based on Transmission Spectra Through

a Polarizer," J. Lightwave Tech 12, to be published (1994).

5. A. J. Barlow, "PMD Measurement Using Fourier Analysis: Investigation of Mode Coupling

Behavior in Fibers," submitted to NIST Symp. on Opt. Fib. Measurements, Boulder,

September 13-15, 1994.

TABLE 1. Inter-Laboratory Comparison

Participant Window (nm)

PMD Coefficient (ps/km^'')

Fiber 1 Fiber 2

A

B

C

D

E

F

G

H

1400-1600

1400-1600

1470-1570

1450-1640

1465-1563

1200-1550

1300-1600

1363-1563

.117

.070

.072

.091

.053

.041

.044

.088

.116

.116

.150

.135

.078

.084

.090

.089

Mean

a

.072

.026

.107

.026

TABLE 2. Measurement Uncertainty

a/<AT>

Fiber 1 Fiber 2

Inherent Statistical Uncertainty (Eq. (2))

Inter-Laboratory Measurement Results (Table 1)

173

0.24

0.36

0.19

0.24

Characterization of dispersion compensating fibers

Ashish M. Vengsarkar and Anne E. Miller

AT&T Bell Laboratories, 600 Mountain Avenue, Murray Hill, NJ 07974

I. Introduction

Highly efficient erbium doped fiber amplifiers with broadband amplifying capability in the 1.53-1.56 |im range are now commercially available. The absence of comparable amplifiers in the 1.31 |im window has spurred system

designers to consider the possibility of upgrading earlier generation 1.31 |im systems for use at 1.55 ^un. The

replacement of bit-rate-dependent electronic regenerators by all-optical bit-rate-independent amplifiers can lead to

i) increased span lengths between amplifiers and/or ii) higher data rates. Optical amplifiers can compensate only for the loss in the system; the issue of dispersion, however, remains unresolved since conventional single-mode fibers embedded in the ground have a zero dispersion wavelength of 1.31 |jm and a positive dispersion of - 17 ps/nm-km

at 1.55 Mm. This positive dispersion places a Umit on the maximum achievable bit-rate distance products - for example, an externally modulated system with a dispersion penalty of less than 1 dB can be operated at a data rate of 10 Gb/s over spans shorter than 62 km.' As a consequence of this limitation, any upgrade of a 1.31 nm system

to a high bitrate system operating at 1.55 |im requires dispersion compensating elements.

Bulk-optic devices for dispersion compensation were proposed several years ago;^^ more recently, there has been

a surge of interest in the development of ail-fiber compensators'"'^ that provide large negative dispersions. The

principle of operation of all these devices is fairly simple - at periodic distances in the system one introduces a

dispersion compensator with a dispersion that is equal in magnitude and opposite in sign to that of the fiber span

to be compensated. Practical fiber-based dispersion compensators can be broadly divided into two categories: i) Fundamental mode designs, where an appropriate choice of the refractive index profiles result in the LPq, mode being

used for dispersion compensation,*"'^ and ii) Higher-order mode designs, where a higher-order spatial mode (typically the LPi, mode) operated close to cutoff delivers the high negative dispersion.'"'

In this paper, we will describe in detail the measurement issues that are specific to the complete characterization of

dispersion compensating fibers (DCFs). At first glance, the characterization of DCFs may be perceived to be very

similar to that of conventional communications fibers since it involves the measurement of several standard fiber parameters such as loss, chromatic dispersion, cutoff wavelength, mode-field diameter, polarization mode dispersion

and geometry/index profile. While one may expect no surprises in the measurement of these parameters, some of

these measurands are more important than others. It is the object of this paper to highlight these differences.

The underlying principle of most DCFs is to enhance the negative waveguide-induced dispersion by using high A's (normahzed index differences) in the core and/or by adding multiple cladding layers of varying refractive-indices around the core. The presence of high A's in the core and a trench (negative A) in the inraiediate cladding layer

raises issues of increased losses and sensitivity to bends. Bend losses are particularly important since several kilometers of fiber may be wound on small spools to achieve a compact package. The variation of any of the aforementioned parameters as a function of bend radius thus becomes an important issue. An unusual distribution of modes may be observed if the design is chosen such that higher order modes can copropagate with the fundamental mode. Inadvertent mode mixing should be avoided for errw-free operation of the compensated system.

//. Performance Evaluators

Five parameters that need to be measured in the characterization of DCFs are now described. The discussion is most

applicable to single-mode DCFs although some of these criteria can also be appUed to the higher-order mode

technique.

A. Chromatic Dispersion

The first parameter of relevance is the dispersion of the compensating fiber. Since conventional fiber (e.g., AT&T

175
5D fiber) has a positive dispersion of -17 ps/nm-km at 1.55 pm, the dispersion compensating fiber should have a

negative dispersion at the same wavelength. It is also desirable that the magnitude of the dispersion of the com|?ensating fiber be very large. Large negative dispersions will lead to short lengths of the dispersion compensating

fiber and thus reduce space requirements at the regenerating stations. Since dispersion scales linearly with length, the total dispersion in a compensated link is given by

where D is the dispersion, L is the length, and die subscripts TL, DCF and 5D denote total link, dispersion

compensating fiber and a conventional 5D (AT&T) fiber, respectively. In calculating the total link dispersion typical

values of = 15.6 and 17.3 ps/nm-km at 1530 and 1560 nm, respectively, can be used.

An additional desirable property of the dispersion performance relates to the slope of the dispersion versus wavelength curve. Since the dispersion of a conventional 5D fiber has a positive slope, the DCF dispersion should

have a negative slope, thus enabling dispersion compensation over a broad wavelength range, specially over the 1530

- 1560 nm range of the erbium-doped fiber amplifiers. For optimal broadband compensation, the ratio of the slopes

of the dispersions (denoted by DO of the DCF and 5D should approximately equal the ratio of the dispersions of

the two fibers, i.e..

^DCF _ ^ DCF

'5D D',3D (2)

The degree of proximity of the slope equality in the above expression sets the dispersion penalties on individual

channels in a wavelength-division-multiplexed (WDM) system.

B. Polarization Mode Dispersion

As in conventional fibers, the effect of PMD can be fairly significant in EXUFs specially if long lengths of fibers are

to be compensated. An added aspect of the measurement of PMD for DCFs is its dependence on bend radius since

the bend-induced birefringence and die winding conditions can affect the PMD. It is therefore recommended that PMD be measured on the final spool size that will be incorporated into the package.

C. Losses

Every IXI!F segment introduce a loss which is typically compensated by a booster amplifier. This additional amplifier,

in turn, adds noise to the system and degrades the optical signal-to-ratio (SNR). This SNR degradation is of concern

to system engineers. An efficient DCF should therefore have low losses.

Intrinsic losses in fibers have been shown to depend on the germania content in the core as well as drawconditions. Communications grade fibers have normalized index differences A of approximately 0.2-0.3% and

typical losses of - 0.20 dB/km. It has been observed that as the core A's increase, the losses gradually increase up

to A = 1.0 % and then rise rapidly for higher A's. Since DCFs have high A's ranging from 1 - 3 % the losses can be expected to be higher than communications-grade fibers. A conventional matched-clad fiber design always allows the propagation of the fundamental mode. Multilayer

waveguides, on the other hand, can be designed to make the LPq, mode leaky beyond a given wavelength. This

property of a multiclad fiber to have a nonzero fundamental mode cutoff wavelength (X^,) is used effectively in some

of the designs by operating die DCF close to in order to achieve large negative dispersion. The presence of

multiple layers and the operation close to cutoff raises the issue of bend sensitivity of the fiber. Bend losses in a fiber can be evaluated theoretically for fibers with simple matched-clad refractive-index profiles; for multilayer

waveguides, however, theoretical predictions are not accurate. Also, the analysis of bend losses depends on the bend

176
radius prescribed by the system designers. An empirical method of evaluating the bend performance of the fundamental mode in fibers involves the calculation of a MAC-number,'* defined by

MAC = MFD

(3)

where MFD is the mode-field diameter and X,. is the cutoff wavelength of the next higher order mode. For matchedclad single-mode fibers it has been shown that a MAC-number < 8.5 ensures a bend-loss < 0.1 dB (@ 1.55 [m) for 50 bends over a mandrel diameter of 60 mm. During the design process, one can simply use the qualitative feature

that smaller MAC-numbers imply better bend performance. By experimentally measuring the bend losses in several multiclad fibers, the designer can then determine the upper bound on the MAC-number. When the loss in a DCF

is plotted as a function of wavelength, the bend loss component tends to start rising rapidly at a wavelength which

depends on the MAC-number and the bend radius. One would ideally like to design the fiber such tiiat > 1560

nm so that only the intrinsic loss would affect the figure of merit of the compensator (defined below).

Another loss mechanism introduced in a packaged DCF module is the splice loss that results fi"om mismatched mode

field diameters, a property we consider next

D. Mode Field Diameter

The mode field diameter (MFD) of the DCF is an important measurement parameter for two reasons. First, when

the DCF is fusion-spliced to a conventional 5D fiber, the induced splice loss is direcUy related to the mismatch of

the MFDs. One way to solve the splice-loss problem is to use the thermally expanded core (TEC) technique" to match the mode-field-radii of the two fibers prior to splicing. The second constraint on the MFD of the DCF is governed by the maximum power that can be launched into the DCF without introducing nonlinearities. The specific nonlinearities of concern are self phase modulation (SPM) and stimulated Brillouin scattering (SBS).'^'' The

maximum amount of power that can be transmitted without inducing SPM or backward SBS is directly related to the cross-sectional area (A,ff) of the mode field and hence to the square of the MFD. In order to minimize the deleterious effects of both these nonlinearities, it is desirable to design the DCFs with large MFDs. The MFD also impacts bend losses of the DCF, as seen in the previous subsection.

E. Modal Noise

Multiple layer fiber designs with larger core diameters allow the copropagation of higher order modes (typically, the LP, , and/or LP02 modes) which can lead to mode-mixing. If the dispersion characteristics of the higher order modes

are different from the fundamental mode dispersion values, modal filters may be needed in the DCF module.

Alternatively, methods of launching only the fundamental mode should be available at the input end of the DCF. One

method that can be used for analysis is viewing the far-field at the output end of the DCF and looking for spatial

patterns that indicate the presence of higher order modes.'* A more accurate measurement technique involves the measurement of the system BER performance as a function of different mode-coupling conditions. This measurement

becomes more important for techniques that require intermodal convertors, e.g., die higher-order mode compensation

technique.

///. Compensator Comparison

A figure-of-merit (FOM) for DCFs has been defined in the literature * as the ratio of the dispersion in the fiber to the loss of the fiber, viz..

The FOM, can be used for a direct comparison of different types of DCFs as well as in system designs to evaluate

FOAfj (ps/nm-dB) = (ps/nm-km)

(4)

ttj (dB/km)

177
the decrease in optical signal to noise ratio (SNR) due to the addition of the DCF. An alternate definition, coined

by Poole," is given by

aAdB/km)

FOM^ (dB/km) = -^-Y—- , (5)

where T is the dispersion D^^ (ps/mn-km) divided by 17 ps/nm-km (dispersion of conventional fiber). In Eq. (5), a smaller FOM2 implies a better DCF.

It must be noted that the FOM should not be the sole criterion on which EX^Fs are compared; for example, a DCF

with Dx = - 51 ps/nm-km and a = 0.22 dB/km would have an impressive FOM of 232 ps/nm-dB but for compensating the dispersion in a 360 km span, one would need 120 km of the DCF ! Also, the loss of the EXI!F, a

in Eqs. (4) - (5), should include not only the intrinsic losses of the fiber but also length-normalized values of i) the bend loss which depends on the diameter of the spool on which the fiber is wrapped, ii) splice losses resulting from

mismatched mode-field radii and iii) other insertion losses introduced by mode-conversion devices (if a higher-order

spatial mode propagates in the fiber). Furthermore, Eqs. (4)-(5) do not take into account the nonlinearity-imposedlimitation on the maximum power that can be transmitted through the DCF. A more comprehensive definition for FOMs is clearly needed.

IV. DCF Examples

Examples of multilayered DCFs fabricated at AT&T are now presented,^" with a partial listing of the properties

described in the previous section.

One multilayered waveguide structure is shown in Fig. 1. This design was primarily chosen for the analysis of the bend performance of the fiber. The preform was drawn to three different outer diameters, namely, 108, 116 and 125

|jm. The corresponding core diameters were measured to be 3.4, 3.6 and 3.9 pm, respectively. A spectral loss measurement was made with 200 meters of the fiber wound on a drum of 30 cm diameter. The data (Fig. 2) showed

a steep rise in the loss at a wavelength ( ) dependent on the fiber diameter. An exact fiber profile was used in a theoretical model ^' for the determination of the mode-field diameter (MFD) of the fundamental mode and the cutoff wavelength ( ) of the higher-order mode of the fiber. The ratio of the two parameters provided the MACnumber, which was seen to be directly related to the wavelength at which the steep bend loss edge was observed.

Table I shows some of the properties of the fibers. In order to maintain > 1560, we see that a MAC-number <

6.0 is desirable. Furthermore, if the fiber is to be wrapped on smaller diameter spools, a smaller MAC-number would

be required.

Table !. Bend loss performance of multiclad fibers wound on a 280 mm diameter drum.

Diameter (pm) (nm) X, (nm) MFD (pm) MAC Dx (ps/nm-km)

108 1410 1114 17.62 15.82 - 150

116 1540 1131 7.02 6.21 - 70

125 1640 1420 4.96 3.49 - 30

A fiber preform with a refractive index profile shown in Fig. 3 was drawn to a core diameter of 3.5 jam. The MFD

of the fibers was calculated to be 5.40 im, while the cutoff wavelengths was estimated to be in 1.46 to 1.50 |4m

range, thus resulting in MAC-numbers of ~ 3.70. The dispersion of the fiber was measured to be -93 ps/nm-km (Fig.

4), with a total loss of 0.40 dB/km when the fiber was wound on a drum of 30 cm diameter. A figure of merit of

232 ps/nm-dB was thus obtained. The dependence of the MAC-number on bend radius was studied by wrapping the

178
fiber on spools of varying diameters. The chromatic dispersion and PMD increase with decreasing bend radius, while

the wavelength at which the loss rises rapidly (XJ decreases, as summarized in Table II.

Table II. Performance variation as a function of bend radius.

Bend Radius (mm) PMD (ps/ Vkm) D;^ (ps/nm-km)

140 1560 0.18 - 93

76 1530 0.19 - 94

1

1 ^ 1485 0.27 - 95 ]

V. Conclusion

Recent advances in preform fabrication techniques and fiber designs have allowed researchers to reach FOMs of 250

-300 ps/nm-dB making the use of DCFs very practical from a system viewpoint.^"-^^ The design issues of DCFs have

been studied fairly exhaustively and the accent is now on production, packaging and qualification. A key issue of

DCFs that intertwines all the above is the measurement of the important parameters. While the measurement

techniques are expected to be the same as those for conventional fibers, it is our hope that this brief summary of the

critical parameters and the trade-offs involved will help set the guidelines for the measurement community.

VI. References

1. A. F. Eb^aie, R. E. Wagner, D. A. Atlas, and D. G. Daut, / Lightwave Technol. 6, pp. 704-709 (1988).

2. E. B. Treacy, IEEE J. Quantum Electron. QE-5, p.454 (1969).

3. O. E. Martinez, IEEE J. Quantum Electron. QE-23, p.59 (1987).

4. A. Yariv, D. Fekete, and D. M. Pepper, Opt. Lett. 4, pp. 52-54 (1979).

5. C. D. Poole, K. T. Nelson, J. M. Wiesenfield, and A. R. McCormick, Opt. Utt. 17, pp. 985-987 (1992).

6. A. M. Vengsarkar, W. A. Reed, and C. D. Poole, Opt. Utt. 17, pp. 1503-1505 (1992).

7. C. D. Poole, J. M. Wiesenfield, and D. J. DiGiovanni, IEEE Photon. Technol. Lett. 5, pp. 134-136 (1993).

8. J. M. Dugan, A. J. Price, M. Ramadan, D. L. Wolf, E. F. Murphy, A. J. Antos, D. K. Smith, and D. W. Hall,

Proc.OFC '92, Post DeadUne Paper, PD-14, pp. 367-368 (1992).

9. H. Izadpanah, C. Lin, J. Gimlett, H. Johnson, W. Way, and P. Kaiser, Proc. OFC '92, Post Deadline Paper, PD15, pp. 371-372 (1992).

10. A. V. Belov, S. I. Miroshnichenko, and V. A. Semenov, OFC '93, Paper ThJ2, pp. 203-204 (1993).

11. A. M. Vengsarkar and W. A. Reed, Opt. Utt. 18, pp. 924-926 (1993).

12. A. M. Vengsarkar, A. E. MUler, and W. A. Reed, Proc.OFC '93, Post Deadline Paper, PD13, pp. 56-59 (1993).

13. S. T. Davey, D. L. WUUams, D. M. Spirit, and B. J. AinsUe. Proc. SPIE 1171, pp. 181-190 (1989).

14. S. -V. Chung, Proc. Intemat. Wire & Cable Symp. 37, pp. 704-709 (1988).

15. H. Hanafusa, M. Horiguchi and J. Noda, Electron. Utt. 27, pp. 1968-1969 (1991).

16. G. P. Agrawal, Nonlinear Fiber Optics, Chapter 4, Academic Press, New York (1989).

17. Y. Aoki, K. Tajima, and I. Mito, J. Lightwave Technol. 6, pp. 710-719 (1988).

18. E. Snitzer and H. Osterberg, /. Opt. Soc. Amer 51, pp.499-505 (1961).

19. C. D. Poole, "Dispersion compensation for Ughtwave systems", OFC Tutorial ThJl, Proc. OFC '94, San Jose, February 1994.

20. A. M. Vengsarkar, A. E. Miller, W. A. Reed, M. Haner, A. H. Gnauck, and K. L. Walker, Paper ThK2, Proc. OFC '94, San Jose, February 1994.

21. T. Lenahan, Bell Sys. Tech. J. 62, pp. 2663-2694 (1983).

22. M. Onishi, H. Ishikawa, M. Shigematsu, H. Kanamori, M. Nishimura, Paper ThKl, Proc. OFC '94, San Jose, February 1994.

179
180
Tailoring of Dispersion Compensation Fibers with

High Compensation Ratios up to 30

R. BoneB (TU Dresden), J. Vobian (FTZ Darmstadt), S. Unger, J. Kirchhof (IPHT Jena)

In the last two years, a considerable interest in the development of dispersion compensating

fibers (DCF) has been taken. This is due to the possibility of upgrading the installed systems

from the 1.3 /xm to the 1.55 jum window using equalizing fibers. At the OFC ' 91 and OFC

'92 some papers were presented with various compensation fibers with a compensation ratio

(CR) up to 30. But this ratio has been achieved with a two-mode fiber which is diflficuh to handle and expensive. With pure monomode fibers only a CR of about 13 has been achieved

[1]. Up to now systematical investigations of such DC fibers have not been carried out.

In this paper we will show methods for tailoring DC fibers with a high CR for monomode

fibers with various profiles. We started with an iterative numerical optimization of the

refractive - index profile and radius parameters with a prescribed CR up to 30 or even higher.

Being interested in the second order dispersion compensation, we have to prescribe the

dispersion values in the vicinity of 1.55/xm. It is necessary that the dispersion decreases

monotonically near X=1.55^m. We have shown [2] that it is not trivial to fiilfill these

preconditions. Principally we can investigate any smooth or nonsmooth refractive-index

profile, but, first of all, we have restrained our studies on W , TC (triple clad) and QC - fibers.

The optimization programme minimizes the standard deviation of the actual dispersion-, field radius- and cutoff frequency values from the prescribed parameters for some discrete

wavelengths with the GAUSS-LEVENBERG-MARQUARDT-algorithm [3]. Although this

method is globally convergent, in general, it may happen that no solution exists for the

prescribed parameters. Nevertheless the algorithm finds solutions (relative refractive-index

steps, radius) in dependence on the initial profile, in most cases. We also control the

PETERMANN-I and PETERMANN-II field radii (wf and w^). These values strongly depend on the spectral dispersion fiinction and cannot be defined independently of the dispersion

behaviour.

Our programme is based on a very stable and accurate method for dispersion calculations

without any numerical differentiation [4]. In this way, we obtained a high stability which is absolutely necessary for the unconstrained optimization method due to the operation near the

cutoff frequency of the LPqi- mode.

In this paper we will present some calculated optimized profiles for the compensation ratios

CR=13, 20 , 30, 40, 50 and the corresponding dispersion and field-radii curves in dependence on the wavelength. Furthermore we will discuss the results and show that technological

barriers make it difficult to achieve CR greater than 30, as a consequence of the high

refractive-index steps (> 3%). Moreover, the wf/w^-ratio increases monotonically to 3 . 4

with increasing CR. This ratio is a measure of the sensitivity of the fiber to deformations. The

required accuracy cannot be implemented in practice.

As an example, we present some spectral dispersion- and field radius curves of fibers with QCprofiles (Fig. 1).

181
Figure 1 . demonstrates that the chromatic dispersion decreases monotonically with

increasing wavelength. With increasing CR the ratios wf/w^ spread monotonically.

First experiments were undertaken to implement DC fibers. As an example, we present results

of a triple clad (TC) fiber fabricated by MCVD. Fig. 2 shows the geometry. The core diameter

is 1.68 |j.m, the cladding diameter 125.2 jim. As the consequence of the very small core

diameter, it is not possible to measure a quantitatively correct n(r)-profile using the

conventional RNF-measuring technique. The core level n(r = 0) cannot be resolved and is, therefore, too low. For the n(r)-measurements shown in Fig. 3 a part of the preform was drawn

to a fiber with an outer diameter of 300 \xm.

The optical loss at 1.55 [im wavelength was measured to be 5 dB/km. This value can be

improved to a certain degree, but a natural limit is given by the high dopant concentration.

In Fig. 4 the measured group delay curve and the calculated (5-term Sellmeier fit) dispersion curve are presented.

The characteristic data are: D (1300 nm) = -247.2 ps/km-nm

D (1540 nm) = -166.7 ps/km-nm

D (1550 nm) = -157.6 ps/km nm

?io= 1687.9 nm.

The chromatic dispersion measurements were performed with the aid of a Raman fiber laser

with a mode-locked Nd:YLF laser used as pump source. The peak power of the optical pulses

is up to 4.5 kW (< 17 W cw power), the pulse width is about 35 ps, the output power stability

is better than ± 0,8 %. The spectral range extends from 1 100 to 1800 nm. A large number of

equidistant measuring points (up to 30) can be obtained. By these means a high accuracy, a good reproducibility can be achieved. The standard deviation is comparable with that of

interferometric measuring methods. It is important to determine the zero-dispersion

wavelength Xq of dispersion compensated ultralong hauls for high-bit rate transmission

systems, soliton systems, e.g., as exactly as possible.

References:

[1] A.M.Vengsarkar, A.E. Miller, M. Haner, A.H. Gnauck, W.A. Reed, K.L. Walker:

Fundamental-mode dispersion-conpensating fibers: disign considerations and experiments

Thk2, pp. 225-227, OFC'94, San Jose

[2] BoneB, R., Vobian, J. , Kirchhof, J. and Unger,S.: Tailoring of Dispersion Compensating

Fibers with high Compensation Ratios up to 30. COST- WORKSHOP 94, OPTICAL

TELECOMMUNICATIONS pp 40-41, Nice

[3] Ortega, J. M. , Rheinboldt, C. : Iterative Solution of nonlinear Equations in several

Variables. New York, Academic Press 1970

[4] BoneB, R., Nowak, W. : Lichtwellenleiter mit besonderem Dispersionsverhalten.

Nachrichtentechnik/Elektronik Berlin 42 (1992) 5,8.184-187

182
1400 1430 IJOO

J2i_ (2)

Fig. 1 QC-profiles, chromatic

dispersion and field

radii. CR=13 (l),

CR=20 (2), CR=30 (3)

D(l550nm) = -280, -320,

-480 ps/km-nm

1300 1350 1400 1450 1500

Fig. 1 QC-pr of iles ,, chromatic dispersion and field radii

CR=13 (1), CR=20 (2), CR=30 (3)

D( l550nm)=-280, -320 and -480ps/km nm
REFRRC7IVE INDEX PROFILE

rodius Cum)

Fig. 3 Refractive index profile of the triple clad fiber
Characterization of Single Polarization Fiber

M. J. Messeriy, R. E. Budewitz, and B. K. Nelson

3M Fiber Optics Laboratory

MS260-5B-08

St. Paul, MN 55144-1000

and

R. C. Mikkelson

Department of Physics and Astronomy

Macalester College

1600 Grand Avenue

St. Paul, MN 55105

Introduction

Single polarization (PZ) optical fiber^ ^'^^ like polarization maintaining (PM) fiber, has a

birefringent core that removes the degeneracy between the two polarization states of

the fundamental mode. Unlike PM fiber, PZ fiber polarizes light; typically, the extinction

is better than 40dB after a few meters. The fiber has recently found favor as a pigtail

for electro-optical devices, and as a polarizing element in sensors.

Most of the important characteristics of the PZ fiber are common to telecommunications

fiber, and are measured by the same techniques. Some are unique, though, and

require new or modified methods. In this paper we describe the measurement of

attenuation, cutoff, and width of the operating band. In addition, we compare the

traditional technique of measuring extinction to the method devised by Feth and Blake^

Fundamental Mode Attenuation

We measure the spectral attenuation of a fiber's two principal polarization states with

commercially available equipment (Photon Kinetics FOA2000), modified by placing a Glan-Thompson polarizer between the source and launch objective. The equipment is not intended for polarization measurements, and has (understandably) two

shortcomings. First, the monochromator slightly polarizes light from the source (by

about 5dB). Second, the objectives are not free of strain and tend to mix the

polarization states. We sidestep the first problem by making a cutback measurement

for each polarization (rather than rotating the polarizer to compare the two), and accept

that the second limits the dynamic range of a polarization measurement to roughly

15dB.

Figure 1 shows the attenuation of the fast and slow polarization states of a 4m sample

of fiber, wrapped eight times on a 75mm mandrel. The logarithmic plot shows that the

loss increases exponentially with wavelength. The data in the Figure, while repeatable, are atypical in two ways. First, it is linear over lOOnm; most curves are only linear over about 70nm. Second, the slope of the two lines is nearly the same; fairly often, the

185
slope of one polarization is significantly steeper than that of the other. (But after 1 50

measurements we see no bias -- the fast state cuts off at a steeper rate as often as the

slow.) We attribute deviations from linearity to the equipment shortcomings described

above. We are not sure if the differences in slope are real or a measurement anomaly.

Early attenuation data suggest that the loss per wrap is a better metric than the loss per

unit length. Figure 2 shows the loss of the fast polarization state of a sample wound 4,

8, and 16 times on a 75mm mandrel; the data are normalized by the number of wraps.

Since we hold the wrap diameter at 75mm, a plot of the loss per unit length looks the

same; to distinguish the two, we also vary the wrap diameter. Figure 3 shows similar

plots for the same sample wound 4, 8, and 24 times on 150mm, 75mm, and 25mm

mandrels, respectively. Figure 3a is a plot of the loss per unit length; Figure 3b is a

plot of the loss per wrap. The latter indicates that the loss per wrap, rather than the

loss per length, stays constant.

We thought these results might have been affected by the 10nm FWHM spectral width

of the monochromator or the equipment limitations described above. To check, we

measured the samples of Figure 3 with laser sources and strain-free optics.

Reassuringly, these points fall in line the others; the laser-based measurements are

plotted as +'s.

Cutoff and Operating Width

Cutoff. We define cutoff to be the wavelength at which the attenuation of the fast

polarization state hits IdB per wrap. We choose this number because, for a typical

number of wraps, it falls within the limits of our equipment. The cutoff of the sample in Figure 1 is 860nm, determined by a least-squares fit.

Operating Width. We define the operating width to be the difference between the

0.01 dB per wrap wavelength of the slow polarization and the IdB per wrap wavelength

(cutoff) of the fast, divided by the fast cutoff wavelength. The operating width of the

sample fiber of Figure 1 is 10%.

Extinction

We measure extinction by the method devised by Feth and Blake^ It requires four

power readings, taken with both the polarizer and analyzer parallel or perpendicular to

the guiding axis, and with one parallel and the other perpendicular to it (both

permutations of the latter). The extinction is the ratio of the power read when the two

are parallel to the guiding axis to the power read when they are perpendicular. We find

that it is typically 35dB to 40dB for a broad-band source such as an SLD, and often

exceeds 70dB (measurement limited by power) for a laser source. A more common technique involves exciting the two principal polarizations equally, then

analyzing the light emitted by the fiber to determine the extinction (the "50/50 launch").

186
Even the best strain-free objectives limit this method to 40-45dB, as Figure 4 shows.

The 50/50 launch saturates at roughly the extinction we measure for the pair of

objectives in our setup (43dB), even though the four-point method suggests that some

samples produce much higher values.

Summary

Early data suggest that the best metric of fiber loss is the attenuation per wrap, which

increases exponentially with wavelength. We define cutoff to be the wavelength at

which the fiber's fast polarization state hits 1dB of loss per wrap, and propose a measure of operating width. We also present data from two hundred extinction

measurements, which show that even strain-free lenses limit the traditional technique to

40dB. In contrast, the technique devised by Feth and Blake appears capable of reliably

measuring much higher values.

References

1. J. R. Simpson et al., "A single-polarization fiber," J. Lightwave Techno!., vol. LT-1,

pp. 370-373, 1983.

2. K. Tajima et al., "A new single-polarization optical fiber," J. Lightwave Techno!., vol.

7, pp. 1499-1503, 1989.

3. M. J. Messerly et a!., "A Broad-band single polarization optical fiber," J. Lightwave

Techno!., vol. 9, pp. 817-820, 1991.

4. Y. Kubo, M. Takagi eta!., "Development of an absolute single-polarization optical

fiber," Sumitomo Etectric Tech. Rev., pp. 60-63, 1991.

5. J. Feth and J. Blake, "Characterization of high birefringence fiber for sensor

applications," Tech. Digest - Symposium on Optica! Fiber Meas., NIST Special

Publication 792, 1990, pp. 109-114, 1990.

187
CO

(AO

-Jc

o

<0 Q

1- Q. 3 (0

li. >

oo

o

o - o

o

o

It o

LU o

c

«

>

[lu/ap] ssoi

in

ino

_3

C

O

in CM Q.

il o

o

o

o

t:

o

oo

o

oo>

3; o

c

«>5

c

X c ^ IJJ ©

« E

^ o o

3 _ rags rr O « U- « « S.S

E

oO

o

_ o

o

.. oCM

oCM

[dejM/gp] sso|

[gp] Mounrn os/OS

CD2,

o

o

o

o.

I w

3

O

u.

rsi

«

O) c

il .23

C

0)

o - o

«2

slow - o

00 ° o

a

ttllllQ o 1 tlllll 1 1< • ' 1

E

c

c

>

5

o

o

o

o

(0M

O

-ic

o « o

w & 3 (0 •-2

o

_ o ^ o

o

o

oo

o

o

c

«

>

E

Em

CM

CM<

E

Em

00

E

Eo

[dejM/gp] sso|

[dej/vvsp] sso|

188
New Direct Measurement Technique of Polarization

Dependent Loss with High Resolution and Repeatability

Naojoiki Mekada, Abdallah AL-Hamdan *, Tomohiro Murakami, Masanori Miyoshi

Santec Photonics Labs., Komaki-shi, Aichi-Ken, Japan

*Santec U.S.A Corp. , Holmdel, New Jersey

Abstract

Polarization Dependent Loss (PDL) of optical components is a very important parameter to specify with high resolution and accuracy. In this letter, we describe a new direct measurement

technique especially designed for low PDL optical components and provide high resolution

(0.001 dB) and high repeatability ( < ± 0.003 dB). This technique is not sensitive to the light source intensity fluctuations.

Introduction

Polarization Dependent Loss (PDL) is the maximum insertion loss variation in optical

components as the input light's state of polarization is varied over all polarization states. Accumulation of PDL for various optical components in light-wave systems can produce large

variations in optical power as the state of polarization changes due to environmental conditions.

Therefore, It is very important to quantify PDL with high resolution and accuracy for optical

components.

Conventional PDL measurement techniques, such as the power meter technique, are Hmited

by the residual PDL of the measurement equipment with an uncertainty of more than 0.01 dB[l].

A technique, with an uncertainty of ± 0.001 dB, was proposed utilizing an erbium-doped fiber

(EDF) as a depolarizer to reduce the detector PDL [2]. The main disadvantage of pervious

techniques is that, they require a highly stable light source. In this letter, the proposed

measurement technique is simple direct measurement technique and not sensitive to the light source intensity flucmations [3].

Measurement Technique

The new technique is based on the fact that the maximum and minimum insertion losses are located at orthogonal polarization states and lie at the same vertical trace on the poincare' sphere.

The measurement set up is shown in Fig. 1. The light emitted by the LED light source is filtered

out using an interference filter with a 3 dB bandwidth of 1.0 nm, then travels through the polarizer

located in the automatic polarization controller to generate a Unear polarization.

After the polarizer, the hght travels into the half wave plate (HWP) which rotates 90° with a step of

A(()° (= 1.8°). This will rotate the polarization state around the equator of the poincare' sphere in

0.25 seconds, fig.2 shows the trace generated from rotating the HWP 90° with a step of A(j)0.

The light continues through the quarter wave plate (QWP) in which the polarization state of the

light is transferred into a vertical plan resulting in a vertical trace for a 90° rotation of the HWP, as shown in fig. 3. For every full rotation of the HWP, the QWP will rotate with a step of Aa^ (= 3.6°) covering a

total of 90° and generating N vertical traces to cover all polarization states. Fig. 4 shows the

poincare' sphere with N vertical traces.

189
PDL is measured within 0.25 seconds for each vertical trace resuhing in N values for PDL, then

the worst case PDL is selected.

PDL = Max. { PDLi, PDL2, PDLn}

Since the PDL measurement of each trace takes only 0.25 seconds, this technique is not sensitive

for the light source power fluctuations. An automatic two dimensional calibration, using a short

SM fiber, compensates for the loss variations of the polarization controller. The detector used in

this experiment is selected with a very low PDL value of less than 0.002 dB.

Measurement Results

Fig. 5 shows the PDL test results for a 30 cm SM fiber without disturbing the measurement.

PDL values ranging from 0.001 dB to 0.003 dB, were measured with a high resolution of 0.001

dB and a repeatability of ± 0.001 dB, measurement was repeated 20 times.

Fig. 6 shows the PDL test results for the same fiber while disturbing the measurement, the

repeatability is ± 0.0025 dB, and measurement was repeated 20 times.

Fig. 7 shows the test results for the beveled fiber with 8° angle. The beveled fiber, with 8° angle,

has a theoretical PDL value of about 0.017 dB. The measured average PDL value is 0.0153 dB

with a repeatability of ± 0.0005 dB, measurement was repeated 100 times. The beveled fiber

might be a good candidate standard for PDL measurement.

Fig. 8 shows the PDL test results for an isolator, a PDL value of 0.08 dB was measured. The

graph also shows the loss variations versus the rotation angle of the HWP and the QWP.

Conclusions

A simple direct measurement technique with a high resolution of 0.001 dB and repeatability

of ± 0.003 dB has been introduced. This technique is not sensitive to the light source intensity

fluctuations while covering all polarization states.

References

[11 Bellcore, TA-NWT-001073, no. 2. [2] Bruce Nyman and Gregory Wolter, IEEE Photonics Technology Letters, VOL. 5, NO. 7

July 1993

[3] Santec Photonics Labs, Optical Module Test System for PDL Measurement Technical

Notes , V 2.9.94

190
Automatic

Polarization Controller

LED O o Filter O -o Dur o O Detector PDL < 0.002 dB

Polarizer HWP QWP

Fig.l

PDL Measurement Experimental Set up
20

10

N z20

Repeatability = +0.001

_L J L

.002 .004 OOG

PDL (dB)

.008 .010

Fig. 5

PDL measurement results for SMF(30cm)

(measurement is repeated 20 times w/o disturbing fiber)

20

10 -

N =20

Repeatability = +0.002

.002 OM .006 PDL (dB)

.008 .01

Fig. 6

PDL measurement results for SMF(30cm)

(measurement is repeated 20 times with disturbing fiber)

100

8Q

60

40

20

N =100

Repeatability = ±0.0005

j_

.010 .012 Ol'l .016 PDL (dB)

.018 .020

PDL measurement

Optical Module Test System

OMT-3100

Fig. 7

PDL measurement results for 8° fiber(30cm)

(measurement is repeated 100 times w/o disturbing)

D.U.T. Optical Isolator Wavelength 1550.0 nm

PDL 0.080 dB

HaK wave ptate ( deg.

)

Quarter wavs pfate

(cteg.)

Fig. 8

PDL Measurement results for an Isolator

(Loss variations-vs-HWP and QWP angle)

192
The Statistics of Polarization Dependent Losses

N. Gisin

Group of Applied Physics

University of Geneva, 1211 Geneva ^, Switzerland

Abstract

The statistics of the attenuation of systems containing several components with polarization dependent losses are described.

1 Introduction and Definitions

In systems containing several elements with Polarization Dependent Losses (PDL) connected

by standard optical fibers, the global attenuation and the global PDL may fluctuate. This will afiect the systems performance [1, 2, 3]. Note that the global attenuation is generally not the

sum of the attenuation (in dB) of the elements! Moreover, this is also true for unpolarized light sources. Indeed, the first PDL element will partially polarize the light, hence the attenuation of

the second PDL element depends on the relative orientation of the two PDL elements and on

the polarization rotation produced by the connecting standard fiber. The fact that the attenuation of systems with PDL may fluctuate has clearly an impact on the characterization of such systems. Indeed, if the attenuation measurements show some

variations, it should not all be attributed to the instrument's limitations, but it should, at least in part, be interpreted as due to the physics of PDL.

For the description of polarization states, we shall use the Poincare sphere. Remember that

the degree of polarization is equal to the length of the vector M:

degree of polarization = |M|,

and that the vector M relates simply to the Stokes parameters So, S — {Si, S2, S^):

The transmission coefficient T through an optics with PDL can be characterized by two

numbers between 0 (opaque) and 1 (transparent): the maximum transmission coefficient Tmax

and the minimum transmission coefficient Tmin- The corresponding polarization states Mmax

and Mrnin need not be Hnear, but they their degree of polarization is always maximal: IMmail =

iMninl — 1- Furthermore, they are always orthogonal: they are opposite on the Poincare

sphere: Mmax = —Mmin- Unpolarized light can be considered as an equal mixture of any pair of

orthogonal polarization. In particular unpolarized light can be considered as an equal mixture of

the polarization states corresponding to the maximum and minimum transmission coefficients.

Hence the corresponding transmission coefficient, T^epoh is the average of the maximum and

minimum transmission coefficients:

rp Tfjiax ~l~ Tmin \ J-depol — \^)

193
Note that such a simple relation does not hold if one expresses the transmission coefficients in

dB.

For PDL we propose the following notation and definition:

T - T

~ ~ T -\-T ^

' -'max I -Lmin

Often it will be useful to define a vector f parallel to Mmax and of norm defined as above. The

transmission coefficient for an arbitrary incident polarization state Min is then given by:

T = Tdepolil + MinT) (3)

Sometimes PDL is defined as the difference between Tmax a-nd Tmin expressed in dB [2, 4]:

PDL^B = lOLogi'^) = 10Log{\±^), (4)

J- min ^ i

but, as for Tdepoi-, we shall see that dB is not a convenient "units" for the present investigation.

2 Combination of several components with PDL

Real optical links in LANs incorporate many PDL elements. Their relative orientations can

not be controlled. Moreover they are not fixed, but fluctuate in time due to the effect of the

connecting fibers. A statistical analyses of the attenuation is thus necessary. In this section we

shall consider the mean attenuation, its standard deviation and the majcimum and minimum

values, all expressed in transmission coefficient (number between 0 and 1) and in dB.

The attenuation of the connecting fibers will not be considered separately, but is included

in the mean attenuation Td^p^i of the PDL elements.

The mean attenuation and its maximum and minimum values are easy to express [5]:

M(Ti,,n) = Td,epol\Tdepol2--TdepolN (5)

TmaxL.N — T^max\Tmax2"-TmaxN (6)

and similarly for Tmin\..N- The global PDL Ti..n however is given by a complicated expression

[6], leading also to intractable expressions for the standard deviation of the global attenuation

<7{Ti..n)- Consequently, in the following we shall restrict ourself to the case of small PDL:

r^- <?C 1. This is anyway the case of practical interest in realistic optical networks. Under this

condition, the global PDL Tl.n and the standard deviation of the global attenuation cr(ri j^)

simphfy to: fL.AT = S,=i..Arf,- + 0(r2) (7)

<t(T,„n) = M(Ti..Ar)y ^^^g ^^ r^ + 0{r') (8)

These approximate expressions have been checks by comparing them to "exact" solutions obtained with a Monte-Carlo numerical simulation. The agreement is excellent. Figure 1 illustrates

the above relations, the maximum and minimum transmission coefficients are also displayed. The

mean losses and PDL of each of the 20 elements have been chosen identical, but the orientations are chosen at random.

194
Decibels axe not appropriate units for fluctuating losses, since the Log of a mean differs from

the mean of the Log. However, under the same condition as above of small PDL, Tj <^ 1, the

expression for the mean and standard deviation simplify:

M{-10Log(Ti..N)) = -10Ej=i,.NLog{Tdepoij)+Oir'^) (9)

.(-ioLo,(T,..)) = ^^^m_Ar^ ^ o(r3) (10)

These approximate expressions have also been checks by comparing them to "exact" solutions

obtained with a Monte-Carlo numerical simulation. Again the agreement is again excellent, as

illiistrated on Figure 2.

3 Conclusion

Polarization Dependent Losses influences the system performance and its characterization. The

coherence of the light propagating in the network becomes relevant. Moreover, even neglecting

this effect, simple concepts like the attenuation become more subtle due to the random orientations of the various components with PDL. As a consequence attenuation is a random number.

Expressions for the mean attenuation and its standard deviation have been derived and checked

against numerical simulations.

For small PDL, the global attenuation is well approximated by the sum of the attenuation

(in dB) of the components. The standard deviation, eq. (10), grows essentially linearly with

the number of components with PDL and is proportional to the square of mean PDL of the

components.

4 Acknowledgments

Financial support by the Swiss Telecom is gratefully acknowledged.

References

[1] E. Lichtman, OFC'94 Technica Digest, pp 257-258, 1994.

[2] CD. Poole and T.E. Darcie, J. Lightwave Tech. 11, 1749, 1993.

[3] Burgmeijer J.W. and al., COST 217 interlaboratory comparison of optical measurement on singlemode fibre couplers, Technical Digest, Symposium on Optical Fiber Measurements, p. 7-10, 1990.

[4] B.L. Heffner, Photon. Tech. Lett. 4, 45, 1992.

[5] The depolarization of the light due to Polarization Mode Dispersion (PMD) is assumed negligible,

that is the coherence of the light source is assumed larger than the delay due to PMD.

[6] N. Gisin, Statistics of Polarization Dependent Losses, preprint. University of Geneva, 1994.

195
0 5 JO 15 20

Number of PDL elements

1. Monte-Carlo simulation of the transmission coefficient through 20 components with PDL.

The top and lower curves represent the maximum and minimum transmissions. The central curve is the mean over 30.000 reaJizations, whereas the the last two curves represent the mean plus and minus one standard deviation. The characteristics of the components have

been kept fixed: Tmax = 0.98 = O.ldB, Tmin = 0.79 = IdBJeTdepol = 89 = 0.53dB, T =

0.1 = 0.9dB, but the relative orientations have been cho.sen at random with a uniform

distribution on the sphere.

0 5 10 15 20

Number of PDL elements

2. Monte-Carlo simulation of the transmission coefficient expressed in dB through 20 components with PDL ^ 0.05 = 0A3dB. In order to simplify the figure, the mean loss has

been subtracted. Hence the central curve (mean loss) is displayed close to zero. The two

bottom and top curves represent the standard deviation a-s found with a 30.000 realizations

numerical simulation and with the approximate expression (10).
Effects of fiber nonlinearity and polarization mode dispersion on

long-distance transmission systems with optical in-line amplifiers

Takamasa Imai, Makoto Murakami, and Tetsuo Takahashi

NTT Transmission Systems Laboratories

1-2356 Take, Yokosuka-shi, Kanagawa, Japan

I. Introduction Recent developments in Er-doped optical amplifiers enable very-long

transmission systems to be created without regenerating repeaters^]. [2]. However, in very

long fibers, nonlinear effects, especially four-wave mixing(FWM)[3]-[9] and self-phase

modulation(SPM)n<^]-n3]^ and polarization mode dispersion(PMD)n4]-[i7] become important

and must be taken into account. This is because they can seriously degrade the transmission

performance.

In this paper, we demonstrate the effects due to nonlinearity and PMD, quantitatively.

Countermeasures to these effects and the key points in measuring fiber and system parameters

are also discussed.

II. Transmission performance degradation due to fiber nonlinearity and PMD

The transmission performance of the system shown in Fig. 1(a), is determined by the

received SNR(signal-to-noise ratio)['^]. It is degraded by intensity noise accumulation and

waveform distortion. Fundamentally, the accumulation in intensity noise is caused by the

amplified spontaneous emission(ASE) as shown in Fig. 1(b). Furthermore, FWM between the

signal and ASE induces excess intensity noise.

SPM causes spectral broadening and results in strong waveform distortion in

combination with group-velocity dispersion(GVD) as shown in Fig. 1(c). Polarization mode

dispersion also degrades the waveform in a long-distance transmission system.

Quantitative impact of these effects is discussed later.

A. Noise enhancement due to fiber nonlinearity Much effort has been devoted to

clarify the noise enhancement effect due to FWM by numerical simulations and analytical

methodst^l-t^l These studies indicate that the effect is significant when the signal wavelength is

very near the fiber zero dispersion wavelength or in the anomalous dispersion region.

To design a practical system, it is important to confirm the effect experimentally.

Accordingly, we performed a recirculating fiber loop experiment and confirmed the intensity

noise enhancement due to fiber nonlinearity as function of the GVD valuet'^l As shown in

Fig. 2, when an optical attenuator replaced the fiber loop, the intensity noise after direct

197
detection increased as per the theoretical plot, and well matched the calculated results

considering only fundamental beat noised 8], This result confirms that excess noise is not

created by the attenuator. Also, no significant excess noise was observed after replacing the

optical attenuator by a fiber with the normal dispersion value of -0.26 ps / nm / km. In contrast

to the above results, we clearly observed the emergence of excess intensity noise at distances over 1500 km when using the fiber with the anomalous dispersion value of +0.35 ps / nm / km.

Figure 3 shows the intensity noise dependence on dispersion for distances up to 6000

km. It is clear that the noise increases with dispersion value for all distances in the anomalous

dispersion region. On the other hand, no significant excess noise was observed in the normal

dispersion region, which agrees with the numerical and analytical results. Considering enhanced noise suppression, these results confirm that long distance

transmission systems should operate in the normal dispersion region.

B. Waveform distortion due to SPM and GVD In long-distance transmission

systems, SPM causes spectrum broadening, which enhances the waveform distortion by GVD.

This effect was studied by numerical simulationH U. Also, pulse width broadening due to SPM

and GVD was discussed analytically! '^l. These calculations indicate that the corresponding

demodulated eye opening penalty is approximated by the following equation.

PeycSPMi^^) =^7(1^1^ / Ttc) B^L^yNLfrep.out {\ - Qxpi- aLrep)} / (aLrep) (1)

D is the chromatic dispersion parameter, A the signal wavelength, c the light velocity in

vacuum, L the transmission distance, the nonlinear coefficient, B the bit rate, Lrep the

repeater spacing, and a the fiber's loss coefficient. Coefficient 77, whose value was found to be around 20 from numerical simulationst'^], depends on the transmitted signal pulse shape.

We confirmed eqn. (1) by a 10 Gbit / s signal transmission experiments^]. The

experimental results shown in Fig. 4 indicate that waveform distortion is negligible, even after

3000 km transmission, when the signal wavelength equals the zero dispersion wavelength.

There is some degradation, however, when the signal wavelength is detuned from the zero

dispersion wavelength. Here, we suppressed the noise enhancement due to FWM by carefully

ordering fibers according to the GVD valuest^l. Quantitatively, as shown in Figs. 5 and 6, the

eye opening penalty is proportional to signal wavelength detuning and repeater output power.

The results calculated from eqn. (1), shown in Figs. 5 and 6, well match the experimental

results.The above results are plotted in Fig. 7 and indicate the accumulated dispersion needed to suppress the eye opening penalty to under 1 dB. The required GVD value is inversely

proportional to the square of the total transmission length.

198
C. Waveform distortion due to PMD Polarization mode dispersion also induces

waveform distortion. The eye opening penalty due to PMD was theoretically and

experimentally confirmedn5]-[i7] under the first order approximationt^O], ignoring spectral

characteristics. These studies indicate that the demodulated eye opening penalty due to PMD,

is approximated by the following equation.

Peye, PMD (dB) = C ( 1 - ypMD )rPMD ? (2)

^PMD relative power launched in the two principal states, At is the differential group

delay time between two principal states, and T is the one bit duration. Coefficient C, whose

value was found to be around 50, depends on the transmitted signal pulse shape.

Experimentally, PANDA fibers were used to confirm the relation instead of a transmission linel^'^l Figure 8 shows the observed waveform degradation due to PMD. Figure

9 shows the corresponding eye opening penalty as a function of PMD. The results calculated

from eqn. (2), shown in Fig. 9, well match the experimental results. Because of random polarization mode coupling in transmission fibers, a statistical

analysis of PMD is necessary[2i]-[22]_ gy considering that the probability density of PMD has a Maxwellian distributiont20],[22]^ mean PMD of 5 ps is required for a 10 Gbit / s system in

order to limit the probability of Q penalty exceeding 0.5 dB to less than 10"9.

III. Countermeasures to noise enhancement and waveform distortion

To suppress both the excess intensity noise due to FWM and waveform distortion due

to SPM and GVD, it is useful to arrange the various fiber sections as described below. At the

points in the fiber of high signal power, negative dispersion fibers should be used to suppress

the excess intensity noise due to FWM. By using positive dispersion fibers only where the

signal power is weak, the average dispersion of the transmission line is tuned to be zero; this

eliminates the waveform distortion due to SPM and GVDf^l.

The measured GVD of the total system must be within the required GVD value shown

in Fig. 7. For example, average GVD must be of the order 10"^ ps / nm / km for a 10 Gbit / s, 9000 km transmission system. To suppress the excess intensity noise, care is needed because even a single fiber, usually 10 to 50 km in length, can exhibit swings in dispersion

valuef23],[24] is, while the overall dispersion value may be negative, some parts of the

fiber may have high, positive dispersion values.

To suppress the waveform degradation due to PMD, the mean PMD of the total system

must not exceed the value shown in Section II. To confirm the mean value of the total system,

we must measure the statistical ensembles at the signal wavelength. To obtain the average PMD

of the fiber, the wavelength ensemble of the PMD is usually measured instead of the statistical ensemble because the average over both ensembles coincidet^^l and the wavelength ensemble

199
can be measured in a short time. However, this method is difficult to apply if the system uses

optical in-line amplifiers because the optical bandwidth of the cascaded EDFAs is relatively

narrow, around several nm or less. Thus, the statistical ensembles of the system must be

obtained, and much time is needed to obtain the mean value because of the very slow

fluctuation of PMD in the installed fiber as shown in Fig. 10t26].

IV. Conclusion In conclusion, the effects of SPM, FWM, and PMD on the performance

of a transmission system with optical in-line amplifiers were demonstrated quantitatively. This

enables system design to include countermeasures to these effects. Also, the key points in

measuring fiber and system parameters were also discussed.

Acknowledgments The authors would like to thank Dr. S. Saito and Mr. A. Naka for

their critical discussion on the theoretical analysis. They also thank Dr. H. Ishio, Dr. T. Ito, and Dr. M. Aiki for their encouragement.

References : [1] T. Imai, M. Murakami, Y. Fukada, M. Aiki and T. Ito, Electron. Lett.,

Vol. 28, pp. 1484- 1485, 1992.

[2] H. Taga, N. Edagawa, H. Tanaka, M. Suzuki, S. Yamamoto, H. Wakabayashi, OFC'93,

Paper PD-1, San Jose, 1993.

[3] J. P. Gordon and L. F. Mollenauer, Opt. Lett., Vol. 15, No. 23, pp. 1351 - 1353, 1990.

[41 M. Murakami and S. Saito, IEEE Photonics Technol. Lett., Vol. 4, No. 11, pp. 1269 - 1272, 1992.

[5] D. Marcuse, IEEE J. Lightwave Technol., Vol. 9, No. 3, pp. 356 - 361, 1991.

[6] D. Marcuse, IEEE J. Lightwave Technol., Vol. 9, No. 10, pp. 1330 -1334, 1991.

[7] K. Kikuchi, IEEE Photonics Technol. Lett, Vol. 5, No. 2, pp. 221 - 223, 1993.

[8] A. Naka, S. Saito, OFC94, Paper FCl, San Jose, 1994

[9] S. Saito, ECOC'93, Paper MoC-1.1, Montreux, 1993.

[10] O. Audouin, L. Prigent and J-P. Hamaide, ECOC91, Paper We A6.1, Paris, 1991.

[1 1] A. Naka and S. Saito, J. Lightwave Technol., Vol. 12, No. 2, pp. 280 - 287, 1994.

[12] D. Marcuse, J. Lightwave Technol., Vol. 10, No. 1, pp. 17 - 21, 1992.

[13] T. Imai, M. Murakami, and A. Naka, Electron. Lett., Vol. 29, No. 16, pp. 1409 - 1410,

1993.

[14] R. E. Wagner, A. F. Elrefaie, OFC'88 Paper TuI6, New Orleans, 1988.

[15] C. D. Poole, R. W. Tkach, A. R. Chraplyvy, and D. A. Fishman, IEEE Photonics

Technol. Lett., Vol. 3, No. 1, pp. 68 - 70, 1991.

[16] F. Bruyere and O. Audouin, IEEE Photonics Technol. Lett., Vol. 6, No. 3, pp. 443 - 445, 1994.

[17] T. Takahashi, T. Imai and M. Aiki, Electron. Lett., Vol. 30, No. 4, pp. 348 - 349, 1994.

[18] N. A. Olsson, J. Lightwave Technol., Vol. 7, pp. 1071 - 1082, 1989.

[19] M. Murakami, Y. Fukada, and T. Imai, ECOC'94, to be presented.

[20] G. J. Foshini and C. D. Poole, J. Lightwave Technol., Vol. 9, No. 1 1, pp. 1439 - 1456.

[21] N. Gisin and J. P. Pellaux, Optics Communications, Vol. 89, pp. 316 - 323, 1992.

[22] N. Gisin, R. Passy, J. C. Bishoff and B. Pemy, IEEE Photonics Technol. Lett., Vol. 5,

No. 7, pp. 819 - 821, 1993.

[23] K. Inoue, J. Lightwave Technol., Vol. 10, No. 11, pp. 1553 -1561.

[24] M. Ohashi and M. Tateda, Electron. Lett., Vol. 29, No. 5, pp. 426 - 428, 1993.

[25] N. Gisin, Optics Communications, Vol. 86, No. 5, pp. 371 - 373, 1991.

[26] T. Takahashi, T. Imai, and M. Aiki, Electron. Lett., Vol. 29, No. 18, pp. 1605 - 1606.

200
Optical

Fiber

Transmitter

Optical

Amplifier

eye opening

optical spectrum

XX

SIGNAL

Frequency

eye opening

optical spectrum

Frequency

(a) System configuration.

ASE noise + Excess noise

due to nonlinearity

I I Excess noise

I /caused by nonlinearity

Accumulated ASE

Frequency

(b) Noise enhancement due to FWM

2S Eye opening degradation

due to GVD or PMD

spectral broadening

due to SPM

Frequency

Frequency Frequency

(c) Eye opening degradation due to SPM, GVD, or PMD.

Fig. 1 Effects due to the fiber nonlinearity and PMD on long-distance transmission systems.

-110

1,000 10,000

distance, km

O Attenuator A D=-0.26ps/nm/km

D=+0.35 ps/nm/km

Fig.2 Measured intensity noise

against the recirculating distance,

(loop length = 30 km,

amplifier output power = 6 dBm)

40 - m 35 - D

30 - >

_J 25 : <a

V)o

20 = z

15 - >

10 -

tr 5

-

I

I A

calculated

values ^ considering

A-==^^:!. .-;::V^:_._Abeat noise.

v7

-0.5 0 0.5 1 Dispersion, ps/nm/km

300 km

1.5

1 ,500 km

3,000 km

6,000 km

Fig. 3 Dispersion dependence of measured

intensity noise relative to the spontaneous

beat noise generated in one loop.

(Broken lines indicate the spontaneous beat noise

calculated for each number of loops.)

201
s

(a) before transmission

(b) after 3,000 km

transmission, A>tSig= Onm

ffl

(c) after 3,000 km

transmission,

AXsig = -0.27 nm

Fig.4 Observed waveform

degradation due to SPM and GVD.

(horizontal: 20 ps / div.)

1 0 Gbit/s

1

3000 km

Repeater spacing = 50 km

: theoretical ,A,0: experimental

Prep.out = 6 dBm

mD

4-

(0c

0)Q.

D)C

'c

<DQ.O

<D>>

LU

5r 1 0 Gbit/s

Repeater spacing = 50 km

: theoretical 0,n : experimental

t^X = -0.57 nm

0 -0.2 -0.4 -0.6 -0.8 -1.0 Signal wavelength detuning from

zero dispersion wavelength, A>.sig nm

Fig. 5 Signal wavelength dependency

of the eye opening penalty.

0,0 1 2 3 4 5

Repeater output signal power, mW

Fig. 6 Repeater output power

dependency of eye opening penalty.

E

1 E

^10-^

COQ.

.9 .2 w 10 O Q.

-5 ^ -3

(0 O 10

O)

CO>

(0 10

Repeater

output power = ^-1 dBm

5 dBm

1 0 Gbit / s, repeater spacing 50 km

1 000 1 0,000

Total transmission length, km

Fig. 7 Required average GVD value to suppress

the eye opening penalty of less than 1 dB

(b) Ar = 54.6 ps

Fig. 8 Observed waveform

degradation due to PMD

(horizontal: 20 ps / div.)

m

•a

c

(UQ.

D)c

'c0

Q.O

0

>>

UJ

10 Gbit/s

A Measured eye

- opening penalty

Oif0 0.1 0.2 0.3 0.4 0.5 PMD of optical transmission line normalized by bit duration

Fig. 9 Power penalty due to PMD

(Solid line indicates calculated results at C = 50)

1 2 3 4 5 6 7

Time, hour

Fig. 1 0 Time evolution of differential group delay

time Ar and power ratio yp^o launched into two input principal states of polarization(Total

fiber cable length = 119 km.)

202
Highly Repeatable Measurement of Nonlinear Refractive Index

by Cross Phase Modulation Method using Depolarized Pump Light

Takatoshi Kate, Yoshiyuki Suetsugu, Masahiro Takagi, Eisuke Sasaoka

and Masayuki Nishimura

Yokohama Research Laboratories, Sumitomo Electric Industries, Ltd.

1 Taya-cho, Sakae-ku, Yokohama 244 JAPAN

Phone No.;+81-45-853-7171

FaxNo.;+81-45-851-1557

High measurement repeatabihty within ±1% has been obtained using depolarized pump

light in measurements of nonlinear refractive index by cross phase modulation (XPM) method.

The standard deviation has been reduced to one sixth compared with the conventional XPM

method. It has been determined that the nonlinear refractive index of dispersion shifted fiber at

1.55 jum was 3.3x10"-° [m-AV] for hnear polarization.

1. Introduction

In recent years, fiber nonlinearitiest^l such as self-phase-modulation (SPM), cross-phasemodulation (XPM), or four wave mixing (FWM), have been drawing much attention in long-haul

transmission systems. Several methods to measure nonlinear refractive index, n2, have been

reported based on those nonlinearitiesf--'^-^], however, measurement repeatability has remained

unimproved. In this paper, method to measure with high repeatability is reported by the XPM

method by depolarizing pump light to avoid uncertainties caused by polarization fluctuation.

2. Measurement of the Nonlinear Refractive Index bv XPM Method

The phase shift O induced by fiber nonhnearities is given by

^=2^i^(Pprobe+2bPpun,p) d)

where Leff is the effective fiber length, A is the vacuum wavelength of the probe light, Agff is the

effective core area, Pprobe is the probe power, Ppump is the pump power, b is the coefficient which

depends on the polarization states of the probe and the pump lights. The first term is the effect of

SPM and the second term is the effect of XPM.

A measurement system by the XPM method is shown in Fig. 1. In this system, probe

signal power is set relatively weak so that no index in test fibers is dominantly caused by amplified

strong pump light through XPM and that the effect of SPM is negligible. When pump light is

203
modulated in its intensity, probe light is modulated in its phase through XPM. Therefore, n2 can be determined by measuring frequency components induced by phase modulation with the self- delayed heterodyne detection technique while changing pump power.

In the conventional no measurement by the XPM method, as well as the SPM method,

repeatability was the major issue. One important point was the polarization dependence on the

phase shift. The maximum phase shift is given at b=l when the pump and the probe lights are

linearly polarized and parallels each other. The minimum phase shift, on the other hand, takes the

value at b=l/3 when they are linearly polarized and become orthogonal. However, because the

polarization states are not maintained in non-polarization maintaining fibers, the observed

maximum phase shift is smaller than the value at b=l, and the observed minimum phase shift is

larger than the value at b=l/3. Therefore n2 in non-polarization maintaining SM fibers had to be

estimated from the average of the maximum and the minimum of the observed phase shifts. And

the other point was the variation of the polarization states of those lights in fibers under test caused

by changes in external conditions. Thus, variances was incurred by the above mentioned factors in

the measurement of n2 by the conventional XPM method.

3. Depolarized Optical Sources

In order to direcdy obtain the average phase shifts and to reduce the variances, the use of

depolarized optical sources, instead of conventional LD, has been proposed as pump waves. Table

1 summarizes polarization extinction ratio in four types of pump sources tested this time. Type D

was the conventional LD, type C was a broadband SLD (Super Luminescence Diode), and types B

and A were a depolarized LD and a depolarized SLD, respectively. The depolarizer consisted of

two polarization maintaining fibers fused with their principal axes rotated by 45 degrees.

Polarization extinction ratio was determined by measuring the maximum and the minimum power

levels transmitted through a polarizer placed at the output while the polarization states of these

pump waves were changed by a polarization controller.

Table 1 Polarization Extinction Ratio for four types of pump sources. Type A B C D*

Construction SLD+depolarizer LD+depolarizer SLD LD

Polarization Extinction Ratio [dB] 0.32 0.70 2.70 12.10

* conventional method

4. Measurement

In order to clearly know the influence of pump polarization extinction ratio, measurements were conducted on rh in a 7 km of polarization-maintaining dispersion shifted fiber (PM-DSF)

with a refractive index difference A = 1.2% and a birefringence of B = 4.61xlCr'^ to exclude

uncertainties of the states of polarization. The maximum and the minimum phase shifts for four

204
types of pump waves were observed while input states of polarization were changed with

polarization controllers. Fig. 2 shows the measured results as a function of polarization extinction

ratio of the pump lights. As the extinction ratio was increased, the maximum and the minimum

phase shifts approached to the ideal ones. And as the pump waves were depolarized to reduce the

extinction ratio, the phase shifts came to converge at the value for random polarization,

independent of the input polarization states. Therefore it was confirmed that the highly repeatable

measurement became possible by sufficiently depolarizing the pump light to remove the

uncertainties of polarization. From this result, a nonlinearity coefficient, n2/Aeff, was obtained as

10.61x10"^° [lAV] corresponding to the average phase shift of 1.75[rad]. Since Aeff of this fiber

was 31.5 [|um2], n2 of the PM-DSF was calculated as 3.34x10"-° [m^AV].

Next no was measured on a 10 km non-polarization-maintaining standard dispersion

shifted fiber (DSF) with A = 1.0% using the four types of pump sources. Because the states of

polarization change in non-polarization-maintaining fibers, the variance of measurement results

caused by imperfect degrees of depolarization is expected to be smaller than the case in PM fibers.

Fig. 3 shows the average and the standard deviation of the measured no/Aeff obtained from five

consecutive measurements on the same fiber while properly changing input polarization states in

each measurement. Measurement conditions like probe power or maximum pump power, etc. were

set nearly equal in all measurements. From Fig. 3, it was found that the average of n2/Aeff obtained

with the four types of pump sources were nearly the same and repeatability was improved by six

times by using the depolarized SLD (±0.6% : [standard deviation]/[averagel) compared with the

conventional method using LD (±3.7%). Obtained n2/Aeff, by using the depolarized SLD,was

6.39xia^° [lAV] with the standard deviation of 0.037x10"^° [lAVJ. Since Aeff of this fiber was

51.4 [jLim^], n2 of the standard DSF was calculated as 3.28x10"-° [m-AV]. The n2 in the DSF was

found smaller than that in PM-DSF because of the smaller refractive index.

5. Conclusion

High repeatability within ±1% was realized on the measurement of nonlinear refractive

index, n2 by XPM method. The use of depolarized SLD having the polarization extinction ratio of

0.32 dB enabled the improvement of the repeatability to be succeeded. The measured n2 for linear

polarization at 1.55 |um was 3.3x10"^° [m^AV] in a standard dispersion shifted fiber.

Reference

[1] G. P. Agrawal, "Nonlinear fiber optics," Academic Press, 1989.

[2] A. Wada, T. O. Tsun and R. Yamauchi, Proceedings of ECOC'92, Berlin, Germany, p45, 1992.

[3] K. S. Kim and R. H. Stolen, Opt. Lett. 19, 259, 1994.

205
Test Fiber

Oscillator Pump

( SLD or LD

J

Self-Delayed Heterodyne

I Detection

Band Pass Filter

Electrical Spectrum

Analyzer

O/E

Fig. 1 Proposed measurement system of nonlinear refractive index by

the XPM method using depolarized pump waves

E

3

E

S Q.

1

-

Maximurnj)_hj_s_e_shift .(ideaj_ casej_b_=JJ_ • SLD+depolarizer

LD+depolarizer

A SLD

° '"^ -n*^ Average.EhasesjTift_Xb^/3)._

Minimum phase shift (ideal case, b=1/3)

.1 1 10 100

Pump Polarization Extinction Ratio [dB]

Fig. 2 Maximum and minimum phase shifts for four types of

pump waves in PM-DSF

o

X

0) 6

<v.

(\J C

• SLD+depolarizer

LD+depolarizer

A SLD

0 LD

directly measured average of the max. and

average the min.

.1

Fig.

1 10

Pump Polarization Extinction Ratio [dB]

100

Average and standard deviation of n2 lA^a for

four types pump waves in standard DSF

206
Automated Fiber Nonlinear Coefficient Measurements using

Self-Phase Modulation Method

Yoshinori NAMIHIRA, Yuichi TANAKA (*),

and Masahito MURE (*),

KDD R&D Laboratories,

2- 1-15, Ohara, Kamifukuoka, Saitama 356 JAPAN

Phone : -h8 1-429-66-7801, Fax : 4-81-429-66-7516

(*) Oyokoden, Lab. Co., Ltd.

3- 1-23, Niizominami, Toda, Saitama 335 JAPAN

Phone : -^81-48-445-6911, Fax : -h8 1-48-445-6901

A bstract : Automated nonlinear coefficient measurement for optical fibers

using self-phase modulation (SPM) method has been developed for the

first time. Nonlinear Kerr refractive index n2 for various dispersion

shifted fibers (DSFs) is found to be around 2.23 x 10-20 m2/W (± 5 %)

for random polarization states at 1.55 |J.m. A comparison of the new

automated measuring SPM method and the conventional SPM method is also presented.

Introduction : Optical nonhnearities limit the maximum power transmitted through

optical fibers and they have become a serious limitation in optically

amplified ultralong transmission spans such as the transoceanic optical

fiber submarine cable systems [1]. The magnitude of the nonlinearity in

optical fiber depends on the nonlinear coefficient of the fiber doping

materials, the power in the optical fibers and the mode confinement.

Recently, the nonlinear coefficient {nilAejf', n2 is the nonlinear Kerr

refaractive index, Aeff is the nonlinear effective area) [2,3] of optical

fibers has been measured by the self-phase modulation (SPM) method

using a 514.5 nm Argon laser in a short fiber (-11 6m) [3], by the SPM

method using a 1319 nm high power mode-locked Nd : YAG Laser in

short optical fibers (-250 m) [4,5], by the cross phase modulation

(XPM) method using a 1553 nm DFB laser in the long (-20 km) optical

fibers [6], by the self-compensated interferometer (SCI) method using a 1064 nm Nd : YAG laser in Er-doped fibers (EDFs) [7], and by the

four wave mixing (FWM) method using two 1555 nm DFB lasers [8].

In this paper, the automated fiber nonlinear coefficient measurements

using the SPM method at 1 .55 jam for various dispersion shifted fibers

(DSFs) are first presented.

Experiment : The experimental setup for automated fiber nonlinear coefficient

measurement using the SPM method at 1.55 jim is shown in Fig.l. The

picosecond transform-Hmited optical pulse was generated using a tunable optical band pass filter and a DFB laser using the gain switching

207
technique [9] with a 100 MHz repetition rate. The input optical pulse

width (FWHM) At was measured by a streak camera with time

resolution of 1 ps and it was around 34 ps [10]. The output spectra from

the DSF at different powers displayed on an oscilloscope as shown in

Fig.2 [10]. The input frequency spectrum was around 13 GHz, which

was measured by a fiber type Fabry-Perot interferometer.

Accordingly, the product of the pulse width Ax and the frequency

spectral width Av is about Ax Av = 0.442, which means that the pulse is

nearly a transform-limited Gaussian pulse [9,10]. In Fig.2, the

frequency spectra are labeled by the maximum phase shift (|)max which

is proportional to the peak power Pq [10]. The output frequency

spectrum at (j)max = 5.5 n has a slightly asymmetric pulse shape. The

main reason is considered to be due to pulse broadening from the

chromatic dispersion and polarisation effects such as polarisation mode

dispersion.

The maximum phase shift (l)max is given by [2,3] ; _ 2k

r max

X

. ^''o (1)

where, LeffSind X are the effective fiber length and wavelength in

vacuum. In the conventional manual method, the nonhnear coefficient

{n2/Aeff) can be obtained from the number of peaks M in the SPMbroardened spectrum and is given approximately by the following

relation ())max = (M - 1/2)7C [2,3]. In the new automated system, the

fiber nonlinear coefficient (n2/Aeff) is obtained automatically from the

relationship between output optical power and peak power trace as shown in Fig. 3. The maximum phase shift ((t)max ) as a function of peak

powers is shown in Fig.4.

Results & Discussion : The nonlinear measurement results of {n2/Aeff) and the estimated

value n2 are summarized in Table 1. In Table 1, the fibers DSFs-1 ~ 4

are commercial DSFs from various fiber manufactures. All the DSFs

have Germanium doped core and silica cladding.

The nonlinear Kerr refractive index n2 can be estimated from the

(n2/Aeff) by multiplying the Aejf. The Aeff is given by [2,3]

2

I \E{r)\ rdr

(2)

The Aeff can be obtained by the computer analysis from the

measured refractive index profiles of the optical fibers [2,3,5,10,11].

However, recently the Aeff is approximately expressed as Aeff =

0.944(71/4) MFD2 (Namihira's relation) [11].

From Table 1 , it was found that the averaged value of n2 of the

conventional manual measurement and automated measurement were

2.28 X 10 -20 m2AV and 2.23 x 10 -20 m2AV, respectively. The 2.23 x

A,ff=2n

208
From Table 1, it was found that the averaged value of n2 of the

conventional manual measurement and automated measurement were

2.28 X 10 -20 m2AV and 2.23 x 10 -20 m2AV, respectively. The 2.23 x 10 -20 m2/W value of ni is in good agreement with that of 2.25 x 10 -20

m2/W value obtained by four wave mixing (FWM) method [8], and also

which lies within 2.1 - 2.6 x 10 "20 m^fW range of recently reported

results of ni values [3,4].

Conclusion : An automated fiber nonlinear coefficient measurement system has

been successfully developed. The nonlinear Kerr refractive index n2 for

DSFs was around 2.23 x 10 -20 m2/W (± 5 %) for random polarization

states at 1.55 jim. The value of n2 was in good agreement with that of

the 2.1 - 2.6 X 10 -20 m2/W range of recently published results.

Acknowledgment : The authors would like to thank Dr. Y. Urano, Dr. T. Yamamoto,

Dr. K. Sakai, Dr. S. Akiba of KDD R&D Labs., and T. Yasumatsu and

1. Kasahara of Oyokoden Lab. Co. Ltd., for their encouragement. We

wish to thank A. Miyata of Univ. of Electro-Communications, and N.

Tanahashi of Nihon Univ. for their experimental support.

References ' [I] D. Marcuse et al. :JLT,vol. 9, LT-9, pp.121-128, 1991.

[2] G. P. Agrawal :"Nonlinear Fiber Optics", Academic Press, 1989.

R. H. Stolen et al.: Phys. Rev. A14, pp. 1448-1453, 1978.

R. H. Stolenet al.,: NIST, Boulder, USA, pp.7 1-75, 1992.

K. S. Kim et al., : Optics Lett., vol.19, no. 14, pp.257-259, 1994.

A. Wadaet al.,: ECOC92, Berlin, Germany, p.42, 1992.

J. E. Meier et al.,:OFMC'93, Sept.,Torino, Italy, pp. 125-128, 1993.

L. Prigentet al., :. IEEE, PTL, vol.5, no. 9, pp.1092-1095, 1993.

M. Nakazawa et al.,: Optics Lett., vol.15, no.l2, pp.715-717, 1990.

[10] Y. Namihira et al.,: OEC94, 14 B2-2, July, Makuhari, Japan, 1994.

[II] Y. Namihira :Electron. Lett., vol.30, no.3, pp.262-264, 1994.

[3]

[4]

[5]

[6]

[7]

[8]

[9]

Optical BandPass

Filter AX = 0.1 nm

Optical BandPass

Filter AX= 1.0 nm

Optical

4^ Attenuator u

Gain-SW DFB-LD

AX = 1.5476 |am At = 30 ps

Optical

BandPass Filter AX = 0.1 nm

Optical PowerMeter

Optical Sampling

Oscilloscope

4^X

Computer Optical Wavelength PowerMeter Meter

5.5k

^ 4.571

^ 3.5k

2.5k

1.5k

Input

Spectrum

0

Fig.l Automated optical fiber nonlinear coefficient

measuring system using SPM method

at 1.55 |im 209

A v = 13GHz

lOOGHz

Frequency Spectra

Fig.2 Examples of frequency

spectra
0.02

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 Peak Power (arbit. unit.)

Fig.3 Example of automated SPM nonlinear coefficient

measurement result

25

20

15

-9- 10

0

0.0 0.2 0.4 0.6 0.8 1.0 1.2 Peak Power [W]

Fig.4 Maximum phase shift ((j) max) as a function of peak powers.

Table 1 Comparison of nonlinear coefficient measurements

from a manual system with the automated system

Fibers

Manual System Automated System

n2/Aeff

-10 [xlO /W]

n2

-20 2 [xlO m AV]

n2/Aeff

[xlO'^^AV]

n2

-20 2^

[xlO m AV]

DSF-1 4.561 2.398 4.268 2.259

DSF-2 4.240 2.245 4.102 2.167

DSF-3 5.280 2.468 5.280 2.468

DSF-4 3.804 2.000 3.880 2.030

Average 4.471 2.278 4.383 2.231

210
High-accuracy Brillouin Gain Spectrum

Measurements of Single-Mode Fibers

Luc THEVENAZ, Marc NIKLES. Philippe ROBERT

EPFL Swiss Federal Institute of Technology, Metrology Lab.

CH-1015 Lausanne, SWITZERLAND

The importance of the measurements of the Brillouin gain spectrum has been growing in the past few

years owing to its potential use for monitoring strains experienced by optical fibers [1] and for

distributed temperature sensing [2]. It is classically measured by launching a highly coherent lightwave

into a test fiber and by observing the amplification of a weak probe signal propagating in the backward

direction [3]. The optical frequency of the probe signal must be stable and tunable to properly scan the

Brillouin gain spectrum. This spectrum lies 12-13 GHz below the pump frequency at a wavelength of

1300 nm and has a Lorenzian line shape with a few tens of MHz FWHM. Generally two different laser sources are used to generate pump and probe signals, although a method using a single laser source, based on spoontaneous Brillouin scattered light analysis has been demonstrated [4]. Frequency

resolution better than 1 MHz is difficult to achieve using all these methods, so that highly accurate measurements are not available so far. Especially only rough estimates for the FWHM linewidth (40-

70 MHz) and the peak gain (1.5-2.5x10-" m/W) have been reported to date [3, 4, 5].

In this contribution, we present a novel method using a single laser source for BGS measurements and

the pump-probe technique. It makes possible reference measurements owing to its ideal inherent

frequency stability and excellent SNR. Systematic measurements of fibers having various core-cladding

index differences were performed, leaving accurate values for Brillouin gain parameters. A general

behavior was observed and relations between Brillouin gain spectrum parameters and fiber core- dadding index difference are confirmed [4].

Optical Isolator

1.32 |im

Nd:YAG

Laser

Coupler p^^p Fibre under

test

Probe

Microwave

Generator

Fig. 1. Experimental setup for Brillouin gain spectrum measurements

The experimental setup is schematically shown in Fig. 1 [6]. Pump and probe signals originate in the same 1.32 ^im Nd:YAG laser and are launched into the test fiber in opposite directions using a coupler.

The isolator blocks the pump light, so that only the probe signal illuminates the detector. The test signal

211
is set at the proper frequency by modulating the laser light using an ultrawideband electro-optic

modulator driven by a synthetised signal generator. This creates two sidebands in the optical spectrum

that are separated from the pump frequency by the modulation frequency. When this frequency

corresponds to the Stokes shift, Brillouin interaction may occur. The lower sideband experiences gain and

grows exponentially as it propagates along the fiber under test, whereas the upper sideband is depleted

and decays exponentially. The sum of this two effects yields an hyperbolic cosine relationship between

amplification and gain coefficient.

l{L) = 2Io COSh{ gB(v)lpLeff]

where lo is the test signal input intensity, gg is the Brillouin gain coefficient, Ip is the pump intensity and

Lgff is the usual effective fiber length for the nonlinear interactions. This relationship is only strictly valid

under the assumption that any effect on the pump signal (gain or depletion) is negligible.

12.78 12.79 12.8 12.81 12.82 12.83 12.84 12.85

Frequency (GHz)

Fig. 2. Measured Brillouin amplification spectrum of a standard single mode fiber (An=5-10'^, core diameter

Slim). Solid line is the numerical fit of the theoretical curve (hyperbolic cosine of a Lorentzian function).

A typical Brillouin gain spectrum measured using this method is shown in Fig. 2. The measured gain curve ^b(v) perfectly fits a Lorentzian function and the Brillouin parameters can therefore be acurately

determined. The length of the different measured fibers was in the 100m-600m range and were wound by

hand to avoid any tension. Regular mechanical winding induces stresses that were observed to bias the

measurements by causing an inhomogeneous broadening of the gain curve. Care was also taken to decrease the test signal amplitude until no effect on the gain curve are measured, meaning a negligible

pump depletion.

Table 1 suminarizes the measurements performed on theses fibers. The Brillouin Stokes shifts range from

11.5 GHz to 12.8 GHz and a clear relationship is observed between Stokes shift and core-cladding index

differences, as shown in Fig. 3., confirming previously reported results [4]. An interesting feature is the

excellent agreement between the extrapolated value for 4n=0 and the Stokes shift calculated for bulk

silica.

Another relationship was found between Stokes shift and Brillouin gain line width, as shown in Fig. 4. The extrapolated value for An=Q is close to the bulk silica value, but the agreement is not perfect. This

may be due to excess phonon damping occurring in the fiber waveguide, though the confidence level of

the measurements made in bulk silica is not known. These measurements provide with a high

212
Cors-cladding An

(X 10-3)

Brillouin shift (GHz)

1 InowlHfh

(MHz) |im

Peak Gain

m/W

Bulk Silica 13.1000 23.0

4.5 12.8527 34.5 5.04

5.0 12.7960 35.7 4.64 1.59 lO-'ll 0.23

5.0 12.8088 35.4 4.82 2.02 10-11 0.28

8.0 12.7191 34.5

10.5 12.3886 40.6

14.0 12.3309 42.5 2.96 1.42 lO-ll 0.52

30.0 11.5094 52.4 1.84 1.91 lO-ll 1.79

Table I. Summary of measurements performed on different fibers, together with aome fibers characteristics and

calculated gain. Values for bulk silica are extrapolated from reported measurements.

confidence level a 35 MHz value for the Brillouin FWHM linewidth of standard telecommunication fibers Un=5-l()-3)atl.32 \im.

Finally the Brillouin gain coefficient go was calculated from the maximum of the BGS after measurement

of pump power and spot size cOq- The effective cross section Aeff was determined using the classical

formula Aeff=n(Oo^. A quite important fluctuation in the gain evaluation (1.4-2 10^^ m/W) is observed but

there is no clear correlation between gain value and core doping. This confirms the assumption that

Brillouin interaction predominantly takes place within the core, so that the effective cross section Agff

must be weighted by the overlap integral between mode field distribution and core index profile [4]. Actually the highest gain is obtained for a fiber with no dip and a very uniform core index distribution,

and lowest gain for a triangular profile. Care was taken to carefully adjust pump and probe

NX

C5

><o

c

3

C

3

O

m

13.2

13 ^

12.8 -j

12.6 -j

12.4 -I 12.2

12

11.8 £

11.6

11.4 -E 11.2

Bulk Silica

-I i_ +

10 20 30

Core-Cladding Index Difference

40

xlO-3

Fig. 3. Relationship between the frequency of the Stokes shift and the core-cladding index differences

213
polarizations, so that polarization scrambling effects are made negligible. This was checked by observing

the absence of amplification in the orthogonal state of polarization.

The difficulty to exactly determine the effective cross section Aeff makes the power gain coefficient golA^ff,

shown in Table I, more suitable for fiber characterisation. Overall amplification can be calculated directly

from pump power and comparative measurements are possible without further calculations using

additional parameters (spot size, index profile).

In conclusion highly accurate measurements of the Brillouin gain curve were performed in fibers with

very different guiding characteristics. These measurements explain the differences observed between

fibers and bulk silica and the scattered value reported by different authors. They also confirmed that the core doping decreases the acoustic velocity - and thus the Stokes shift - and the phonon lifetime, causing

a broader linewidth, and has a negligible effect on the gain coefficient.

Acknowledgments

The authors would like to thank the Swiss PTT, Cables Cortaillod S.A. and PSEL fund for their kind

support and Cabloptic S.A. to have, in addition to its support, supplied the different optical fibers. A

patent is pending for this measuring technique.

References

[1] T. Horigushi, T. Kurashima and M. Tateda, IEEE Phot. Tech. Lett., vol. 2 (1990), no. 5, pp. 352-354.

[2] X. Bao, D. J. Webb and D. A. Jackson, Opt. Lett., Vol. 18 (1993), no. 7, pp. 552-554.

[3] N. Shibata, K. Okamoto and Y. Azuma,, J. Opt. Soc. Am. B, Vol. 6 (1989), no. 6, pp. 1167-1173.

[4] R.W. Tkach, A.R. Chraplyvy, R.M. Derosier, Electron. Lett. Vol. 22 (1986)., no. 19, pp 1011-1013.

[5] T.O. Tsun, A, Wada, T. Sakai, R. Yamauchi, Electron. Lett. Vol. 28 (1992)., no. 3, pp 247-249.

[6] M. Nikles, L. Thevenaz and P. Robert, Proc. Nonlinear Guided-Wave Phenomena (Cambridge

1993), OSA publications, paper MD 5

214
Guided Acoustic-Wave Brillouin Scattering as an Ultrasensitive

Probe of Optical Fiber Geometry

Benjamin I. Greene

AT&T Bell Laboratories, 600 Mountain Ave, Murray Hill, NJ 07974

Peter N. Saeta

NIST/JILA, Campus Box 440, University of Colorado, Boulder, CO 80309-0440

Abstract

Guided acoustic-wave Brillouin scattering measurements were performed on a variety of

different 20-cm lengths of optical fibers. The detailed line shapes of the lowest lying resonance at ca. 22 MHz were observed to vary significantly from sample to sample. We have attributed these

variations to 0.01 - 0.1 /im distributed deviations from a perfect cylindrical geometry.

Introduction

Pioneering work by two groups in the mid 1980's predicted and experimentally verified the

phenomenon of GAWBS (guided acoustic-wave Brillouin scattering) in optical fibers. ^'-^ The

effect has its origin in radial acoustic oscillations which are bounded as a consequence of

cylindrical fiber geometry. With expectations based upon other high-Q resonance phenomena,"^ we

set out to explore the possibility that the GAWBS line shape could be a very sensitive probe of

deviations from ideal geometric form. Below we present high-resolution spectral data on the

lowest lying GAWBS resonance from optical fiber samples. We have observed significant

variations in these line shapes, and speculate how variations in the size or shape of tlbers could be

manifested in the data.

Experimental

GAWBS spectroscopy was performed as originally described by Shelby er al. ^-^ Briefly,

we use a heterodyne technique, relying on the acoustically driven index fluctuations at the core to shift (add sidebands to) the fundamental laser frequency. The sidebands are mixed, via the

nonlinear response of a pholodiode, with the laser fundamental, and the difference frequency is detected with a microwave spectrum analyzer. Data were taken in the depolarized configuration

primarily for convenience. In this configuration only those acoustic modes resulting in

macroscopic birefringence {i.e., non-radially symmetric modes) are observed.

Results

Several fibers were studied, including many commercial products and some "research" fibers.

All samples were required to be single mode at our laser wavelength (900 nm), excluding from

study, therefore, standard production telecommunication fiber. The fiber type from which most

data were collected, including individual samples which gave the narrowest low frequency

resonances, was Corning "Flexcore 1060." Two examples of the lowest- frequency resonance

taken on different samples from the same 100-m Flexcore spool are shown in Fig. 1. The most

common spectral shape, one that occurred in approximately 80% of the Flexcore samples, was

characterized by a single (or nearly single) peak roughly 25-kHz wide, and a broad sideband or shoulder extending by as much as 80 kHz. This band shape is intermediate between those shown

in the figure. Considerably less probable were either a single clean peak with a FWHM as small as

1 1 kHz, or an overlapping superposition of several spectral features having a composite width in excess of 120 kHz. Spectra of a Newport Research Corporation "NRC-82()" fiber were taken (not

shown) which consistently gave relatively broad (~60-kHz) single-peaked spectral signatures.

215
_0.2 ' ' I I ^ I I I I ' I ' ' '—' I I I '—' ' ' ^ ' I ' ' I 22.40 22.45 22.50 22.55 22.60 22.65 22.70

MHz

Figure 1 GAWBS signals from two samples from the same spool of Corning "Flexcore 1060".

The naiTow resonance has a FWHM of 1 1 kHz.

Additionally, lowest-order GAWBS resonances of "research drawn fibers" were observed to frequently exhibit distinct multiple peaks with individual widths on the order of 10 to 15 kHz.

Discussion

A discussion of phenomenological linewidths must first establish those factors entering into

the homogenous width. For the current case, we suggest two significant homogeneous

broadening mechanisms: material damping and transmission loss at the air-glass boundary. The

literature contains several measurements of acoustic loss in silica, the most relevant to our

frequency range by Fraser, Krause, and Meitzler.'^ These authors report a value of the absoiption

coefficient (a) for the transverse acoustic wave at 22 MHz of 2.12 x l()-2 dB/cm. Relating the

resonant linewidth (A, FWHM) to a (the base-e absorption coefficient) as A = a / 7t, where v is the velocity of sound, predicts a homogeneous width nearly 20 times nairower than that actually

obsei^ved.^ We note that optical fibers are not pure silica rods, but more accurately contain a core of doped silica. For our samples, the cores are roughly 5-8 /im in diameter and consist of

germanosilicate glass, doped at not more than 5 mole-% gemianium. Although we could not find

published acoustic loss data on germanosilicate glass, unpublished data for 10 mole-% samples

indicate these materials to be only a factor of 3 more attenuating at 20 MHz than pure silica.^

Taken together with the fact that the core constitutes only a few percent of the acoustic path, it seems impossible to attribute the obsewed 1 1-kHz width (Fig. 1) to acoustic absorption alone.

Incomplete reflection is another loss mechanism. The core-cladding reflective loss is estimated

with the aid of density and velocity data to be minuscule {R = 10~^)J The air-cladding loss,

however, appears to be significant (T = 0.2%, R = 99.8%). Converting this to an effective a

216
yields a value of 0.16 cm~l (arrived at by assuming a 0.2% loss per traverse across the fiber 125-

IJm diameter), resulting in a 20-kHz component to the linewidth. This appears to indicate that in

the narrowest cases, our linewidths are determined by transmission loss at the silica-air boundary.

This conjecture could be verified by performing measurements in vacuo, where the acoustic

reflection at the outer boundary should become unity. We did not attempt this measurement.

We now discuss possible causes of broader and more complex line shapes. Fluctuations in fiber diameter, deviations from perfect circularity of the cross section, and compositional variations

affecting the velocity of sound could all broaden low-frequency GAWBS resonances. The easiest

to appraise of these with respect to effects on the GAWBS linewidth are variations in diameter.

Diameter fluctuations would simply shift the resonant peak position by the same fractional extent as

the dimensional variation. For instance, a 0.1 -//m change in size for a i25.0-/im-diameter fiber

would shift the GAWBS resonance at 22 MHz by 17.6 kHz (see Fig. 1). Diameter fluctuations

distributed along the sample could therefore result in broadened or distorted hne shapes, depending

in detail on the exact distribution of diameters.

Deviations from perfect circularity of the fiber cross section could also perturb GAWBS

resonances. Depolarized GAWBS is produced by vibrational modes depending on az,imuthal angle

0 as either cos 2(f) or sin 20.^-^ In a perfect circular fiber the two modes are degenerate. Slight

distortions of the cross section from circularity perturb the vibrational modes and, in general, lift the two-fold degeneracy of the modes producing depolarized GAWBS light. This problem can, in principle, be solved in perturbation theory by treating the distortion as a small perturbation of the

solution for the perfect circular case. We have not carried out such a procedure. However, the

simpler problem of the vibrational modes of a nearly circular membrane is discussed by Nayfeh.^

In that case, a distortion of the circumference of the form sin 2m0 or cos 2w0 lifts the degeneracy

between the sin m0 and cos ni(p modes {i.e., of half the period). Furthermore, if e is the relative

amplitude of the distortion to the radius, the two eigenfrequencies are shifted by ±£0), where d) is the eigenfrequency in the absence of distortion. Just as for variations in the size of the fiber, the

relative frequency shift is equal to the relative amplitude of the variation. Assuming that a similar

result holds for the vibrations in the fiber cross section, a splitting between the depolarized

GAWBS modes would arise from perturbations of the form cos 40 and sin 40 in the fiber cross

section. These correspond to a slight "squaring" of the circle. Geometric distortions in optical fibers have been documented and studied by analysis of

backscattered light arising from a laser beam that is incident at right angles to the fiber axis.^^-'^

These investigations leave little doubt that microscopic geometric deviations can exist in optical

fibers. Finally, interferometric micrographs indicate that both radially and longitudinally periodic

fluctuations in refractive index, of order 5x10-^, can exist in prefonns and drawn optical fibers.

The details of these latter effects are somewhat specific to the fabrication technique (MOCVD or VAD). The index, density, or compositional fluctuations are observed to be periodic on a scale

small compared to acoustic wavelengths. Furthermore, observed modulations appear to be radially

symmetric. Therefore, their impact on the GAWBS linewidth would be expected to be minimal.

Work on bulk samples coirelating the effects of thennal processing on the acoustic properties

of vitreous Si02 has been presented. ^''•''^ In the context of fiber process technology, one might

envision pathological cases where non-radially symmetric compositional inhomogeneities could

contribute to the inhomogeneous GAWBS linewidth. Lacking any specific data along these lines,

we believe that the geometric variations discussed above are the most likely explanation for our

current data.

Conclusion

In conclusion, we have investigated in detail the line shape of the lowest GAWBS resonance

in nominal 125-/im-diameter optical fibers. Spectra typically displayed significant amounts of

217
inhomogeneous broadening. We attribute these effects to distributed variations in the fibers' outer

dimensions. For measurements performed in air, we estimate a homogeneous GAWBS linewidth

at 22 MHz of -10 kHz, and ascribe this width to incomplete acoustic reflection at the fiber-air

boundary. This technique is therefore expected to be sensitive to radial geometric distortions on

the 0.01-/im length scale. Our measurements indicate that 0.01 to 0.1-^m deviations from perfect

cylindrical geometry are statistically very common in commercial optical fibers.

Acknowledgment

The authors wish to thank J. T. Krause for sharing unpublished acoustic attenuation data on

silicate glasses with us, and D. J. DiGiovanni for providing samples of limited-production research

fibers.

References

' P. J. Thomas, N. L. Rowell, H. M. van Driel, and G. I. Stegeman, Phys. Rev. B 19, 4986

(1979).

2 R. M. Shelby, M. D. Levenson, and P. W. Bayer, Phys. Rev. B 31, 5244 (1985).

3 R. M. Shelby, M. D. Levenson, and P. W. Bayer, Phys. Rev. Lett. 54, 939 (1985).

^ S. C. Hill and R. E. Benner, J. Opt. Soc. Am. B 3, 1509 (1986).

^ D. B. Fraser, J. T. Krause, and A. H. Meitzler, Appl. Phys. Lett. 11, 308 (1967).

^ G. E. Durand and A. S. Pine, J. Quantum Electron. QE-4, 523 (1968).

^ J. T. Krause, unpublished data.

^ R. N. Thurston, J. Acoust. Soc. Am. 64, 1 (1978).

^ A. H. Nayfeh, Introduction to Perturbation Techniques (Wiley, New York, 1981).

'0 H. M. Presby, Appl. Opt. 15, 492 (1976).

11 H. M. Presby, Appl. Opt. 16, 695 (1977).

1^ T. Izawa and S. Sudo, Optical Fibers: Materials and Fabrication (D. Reidel, Kluwer

Academic PubHshers Group, Noi-well, Massachusetts, 1987).

1-^ J. T. Krause, J. Appl. Phys. 42, 3035 (1971).

14 D. B. Fraser, J. Appl. Phys. 39, 5868 (1968).

218
Albert, J. 85 AL-Hamdan, A. 189

Artiglia, M. 25, 61, 69

Aso, 0. 159

Atkins, R.M. 81

Baines, J. 125 Baney, D.M. 91 Bao, Y. 97 Barlow, A.J. 167, 171 Becher, R.A. 117

Bilodeau, F. 85 BoneB, R. 181 Brehm, C. 133 Budewitz, R.E. 185

Bures, J. 101

Caviglia, F. 49

Chiantore, A. 61, 69

Chowdhury, D.Q. 155, 163

Cottino, E. 25

Daxhelet, X. 101 De Bernardi, C. 113 Di Vita, P. 69

DiGiovanni, D.J. 141 Drapela, T.J. 129 Ducos, L. 45

Emig, K.A. 171 Enard, A. 113

Facq, P. 45

Ferri, G. 61, 69

Fleming, S.C. 15 Franzen, D.L. 129

Frehill, C. 125

Furuicawa, S. 39

Gardner, W.B. 171

Gini, E. 113

Gisin, N. 149, 193 Godbout, N. 101 Goldstein, E.L. 19 Greene, B.I. 141, 215

Grindstaff, S. 171

Haigh, N.R. 77

Harris, D.O. 35

Hattori, Y. 145 Henry, C.H. 105

Hentschel, C. 7 Hernday, P. 171

Hill, K.O. 85 Horiguchi, M. 109

Horiguchi, T. 73

AUTHOR INDEX

Hsu, K. 97 Hugenberg, K.F. 117, 121

Imai, T. 197

Ishida, H. 145

Ito, S. 57

Izumita, H. 39

Johnson, D.C. 85

Jones, T.C.E. 77

Jones, W.C. 171

Kato, T. 203

Kawanishi, S. 31 Kirchof, J. 181 Kitagawa, T. 85 Koyamada, Y. 39, 73

Kiing, A. 113 Kurashima, T. 73

Lacroix, S. 101 Leckel, E. 7

Li, M.J. 53

Lipinski, G. 133 Lowry, M. 117

Malo, B. 85 McConaghy, C.F. 117, 121 Mekada, N. 189 Mercereau, 0. 133 Messerly, M.J. 185 Mikkelson, R.C. 185

Miller, A.E. 141, 175

Miller, CM. 97

Miyoshi, M. 189 Morasca, S. 113 Mori, K. 31 Morioka, T. 31 Murakami, M. 197 Murakami, T. 189 Mure, M. 207

Namihira, Y. 207

Nebout, P. 101 Nelson, B.K. 185 Newton, S.A. 91

Nikles, M. 211

Nishimura, M. 203

Noell, W. 113 Nolan, D.A. 155, 163

Ogawa, S. 145 Ogoshi, H. 159 Ohshima, I. 159

Olson, T. 171 Ozawa, K. 145

Poole, S.B. 1 Pozzi, F. 113

Raine, K. 125

Ricaldone, P. 49

Robert, P. 211

Roland, D.K. 171 Rossaro, A. 61

Saeta, P.N. 141, 215

Sankawa, I. 39

Saravanos, C. 53

Sargis, P.D. 121

Saruwatari, M. 31 Sasaoka, E. 203

Smithgall, D.H. 137

Sorin, W.V. 91 Spiekman, L.H. 113

Stocklein, W. 65

Suetsugu, Y. 203

Sweider, D. 117

Takada, K. 109

Takagi, M. 203

Takahashi, T. 197

Takara, H. 31 Tanaka, Y. 207

Tanji, H. 57

Theriault, S. 85

Thevenaz, L. 113, 211 Throckmorton, R.A. 35 Townley-Smith, P. 53

Unger, C. 65

Unger, S. 181

Vailleau, G.P. 133 van Brug, H. 113 van der Plaats, J.C. 7

Vengsarkar, A.E. 175

Vinchant, J.-F. 113 Vobian, J. 181 Vodjdani, N. 113

Warder, J. 53 Weiershausen, W. 113 Willems, F.W. 7, 15

Yaffe, H.H. 105

Yamada, H. 109

Young, M. 129

Zengerle, R. 113

• U.S. GOVEKKMENT PRINTING OFFICE: lP94-3nn-568/n303A

The National Institute of Standards and Technology was established in 1988 by Congress to "assist

industry in the development of technology . . . needed to innprove product quality, to modernize

manufacturing processes, to ensure product reliability . . . and to facilitate rapid commercialization . . . of products based on new scientific discoveries."

NIST, originally founded as the National Bureau of Standards in 1901, works to strengthen U.S.

industry's competitiveness; advance science and engineering; and improve public health, safety, and the

environment. One of the agency's basic functions is to develop, maintain, and retain custody of the

national standards of measurement, and provide the means and methods for comparing standards used

in science, engineering, manufacturing, commerce, industry, and education with the standards adopted

or recognized by the Federal Government.

As an agency of the U.S. Commerce Department's Technology Administration, NIST conducts basic

and applied research in the physical sciences and engineering and performs related services. The Institute does generic and precompetitive work on new and advanced technologies. NIST's research facilities are located at Gaithersburg, MD 20899, and at Boulder, CO 80303. Major technical operating units and their principal activities are listed below. For more information contact the Public Inquiries Desk, 301-975-3058.

Technology Services

• Manufacturing Technology Centers Program

• Standards Services

• Technology Commercialization

• Measurement Services

• Technology Evaluation and Assessment

• Information Services

Electronics and Electrical Engineering

Laboratory

• Microelectronics

• Law Enforcement Standards

• Electricity

• Semiconductor Electronics

• Electromagnetic Fields'

• Electromagnetic Technology^

Chemical Science and Technology

Laboratory

• Biotechnology

• Chemical Engineering'

• Chemical Kinetics and Thermodynamics

• Inorganic Analytical Research

• Organic Analytical Research

• Process Measurements

• Surface and Microanalysis Science

• Thermophysics^

Physics Laboratory

• Electron and Optical Physics

• Atomic Physics

• Molecular Physics

• Radiometric Physics

• Quantum Metrology

• Ionizing Radiation

• Time and Frequency'

• Quantum Physics'

Manufacturing Engineering Laboratory

• Precision Engineering

• Automated Production Technology

• Robot Systems

• Factory Automation

• Fabrication Technology

Materials Science and Engineering

Laboratory

• Intelligent Processing of Materials

• Ceramics

• Materials Reliability'

• Polymers

• Metallurgy

• Reactor Radiation

Building and Fire Research Laboratory

• Structures

• Building Materials

• Building Environment

• Fire Science and Engineering

• Fire Measurement and Research

Computer Systems Laboratory

• Information Systems Engineering

• Systems and Software Technology

• Computer Security

• Systems and Network Architecture

• Advanced Systems

Computing and Applied Mathematics

Laboratory

• Applied and Computational Mathematics^

• Statistical Engineering^

• Scientific Computing Environments^

• Computer Services^

• Computer Systems and Communications^

• Information Systems

'At Boulder, CO 80303. ^Some elements at Boulder, CO 80303.
U.S. Department of Commerce

Technology Administration

National Institute of Standards and Technology

325 Broadway

Boulder, Colorado 80303-3328

OFFICIAL BUSINESS

PENALTY FOR PRIVATE USE, $300